This paper deals with the mechanical behavior of hollow concrete cylinders prestressed with nickel-titanium (Ni-Ti)-shape memory alloy (SMA) wires wound around them. Prestresses can be created by the thermal activation of the memory effect of SMA wire placed on the outer surface of concrete cylinders. In order to assess the stress level in concrete, a model was used to analyze the thermal stresses in the concrete shell resulting from a temperature gradient in the thickness. Another model was used to calculate the circular concentric loading applied by the wound wire resulting from the impairment of its memory effect by the concrete cylinder. Finally, longitudinal compression tests were performed on the prestressed hollow cylinders. Longitudinal and circumferential strains were measured using gauges located on the surfaces of the hollow cylinders. The tests were performed almost one year after the application of prestressing by means of Ni-Ti SMA wire, confirming that the residual stress in the wire remained present. It may therefore be concluded that the prestressing of concrete elements designed with the use of Ni-Ti SMA material is effective for a long time. The present paper is a follow-up of our previous study dedicated to the prestressing process of hollow concrete cylinders by means of nickel-titanium (Ni-Ti)-shape memory alloy (SMA) wires wound around them . The sitIn our earlier work , sectionthe first objective is to explain the occurrence of failure and to discuss safety in the strengthening of cylindrical elements with the use of SMA wires;the second objective is to present the new results of longitudinal compression tests of hollow cylinders prestressed with SMA wires. The influence of prestressing on the structural behavior of the shell constitutes a fundamental piece of information for the potential applications of the prestressing method.The field of stress in concrete is created as a result of the thermal loading of the cylinder (as a consequence of the heating of the SMA wire placed on its outer surface and the heat diffusion to the concrete) and as a result of the circular concentric loading applied by the wound wires (due to the impaired shape memory effect of the SMA wire). It can be assumed that the failure of the tested cylinders was the consequence of the simultaneous action of both loads. This paper focuses primarily on the analysis of the state of stress in the two destroyed cylinders. To this end, the purpose of the paper is two-fold:For both of these objectives, a key point is that the concrete cylinders are hollow, which is different from previous studies. From a measurement point of view, this allowed us to measure the strain on the internal surface of the cylinders. From an application point of view, we propose to extend the use of SMAs to such hollow components once safety issues are managed.An additional objective of the paper is related to the general durability of prestressing based on coiled SMA wires. In our previous work , the evaa theoretical analysis of thermal strains and stresses in the hollow cylinder shells included in the research program described in our previous research ;a theoretical analysis of hollow cylinders subjected to the mechanical action of a given number of winds of SMA wire for all the cylinders tested in our previous research ;experiments on hollow cylinders subjected to compression forces applied at their end surfaces for the specimens available for the compression tests  to the top (x = H) of the cylinder shell.The calculation of thermal stresses due to a thermal gradient Δd Elliot  and Serud Elliot , based oBending moment in the longitudinal direction :Bending moment in the circumferential direction :Force in circumferential direction :xσ and ϕσ, respectively) on the wall surfaces of the concrete, where:indices “i” and “o” refer to the inner and outer surface of the hollow cylinder, respectively;cW is the section strength index, cA is the area of hollow cylinder cross section.Finally, the fundamental equations relating to the strength of the materials under local stress in the form used in , writtenStress in the circumferential direction:Stress in the longitudinal direction:T in the wall thickness. In all the calculations and diagrams in this work, the conventional signs used for stress and strain in concrete elements analysis is used: “+” for compression and “−“ for tension. For each type of hollow cylinder, the stress values in the longitudinal and circumferential directions were calculated for the assumed ΔT values. It is worth underlining that the relation between the stress values and the thermal gradients is linear. Another important statement is that tensile stress of concrete surfaces is always generated on the opposite side to the side subjected to higher temperatures. For instance, stress distributions in the longitudinal and circumferential directions on both surfaces of the cylinder with a 200 mm external diameter for several of the temperature gradients applied are presented in fct = −4.4 MPa in axial loading is also displayed in the graphs. It can be noted that negative temperature gradients are possible, for instance, during the cooling time of the cylinders, when the inner surface cools more slowly than the outer surface, which is exposed to the ambient air flow.The calculation was performed for each type of hollow cylinder while accounting for the relevant temperature gradients ΔT = 20 °C. The tensile stress in the concrete reaches a value close to the tensile strength of concrete (−4.4 MPa). Compression stresses in the concrete do not exceed 5 MPa, while the effective compression strength of the concrete is 46.4 MPa. The tensile stress distribution in the longitudinal direction of the cylinder is nearly linear in the length section between 0.10 m and 0.40 m.On the basis of the analysis of the calculated stress values for the three diameters , it was concluded that the destruction of the hollow cylinder under the action of the thermal gradient may only be observed at a surface temperature difference of ΔT = To − Ti = −14.2 °C  must now be considered.ctmf, cracking on the internal surface of the hollow cylinder may be created over the whole internal circumference. The evolution of the crack width, together with its depth in the hollow cylinder wall, are functions of the prestressing force in the wire turns. In the theoretical calculations, a constant value of prestressing force in every turn was assumed. For the hollow cylinders with diameters of 200 mm, 250 mm and 300 mm, prestressing force values in the range of 150 N to 1500 N, 217 N to 1420 N and 235 N to 1740 N were assumed, respectively, in accordance with the results reported in our previous paper [Following the prescriptions of the presented testing procedure, all the tested hollow cylinders included in the research program described in our previous paper  were subStatic analysis of the hollow cylinders was performed with the use of NOAM calculation software by J. Michno ,17. The The local stress values may be calculated with the analytical model formulated for the cylindrical shell. The distributions of the circumferential and longitudinal stress and shear forces in the wall of the analyzed cylinders are shown in The distribution of the stress calculated for the cylindrical wall of the cylinder depends on the number of wire windings and on the value of the force activated in the SMA wire. In all of the calculations, a constant force in the wire was assumed. The force values are included in In fct = −4.4 MPa was generated under the prestressing force at 807 N, and when the force reached the value of 1420 N, the tensile strength was exceeded by 80 %  and oneAt this point, six electric resistance wire strain gauges were also installed on the internal surface of each hollow cylinder, as described below . Three of these gauges were located in the circumferential direction and another three in the longitudinal direction of the cylinder. The temperature of the internal surface of the hollow cylinders was also measured.oP = 640 N, termed R1;hollow cylinder with wound wire of prestressing force of oP = 1280 N, termed R2;hollow cylinder with wound wire of prestressing force of reference hollow cylinder—no wire, termed R3.The experimental tests on the hollow cylinders were performed with the use of the ZWICK 1600 type multifunctional testing device. The tests were performed with the deformation control and the deformation increase per unit time was kept constant at 0.05 mm/min. During the test, the value of the applied force was recorded. It was planned that the test would be stopped when the measured force dropped by 20% compared to the maximal value in the test. As described above, three concrete hollow cylinder specimens were subjected to testing two of which were prestressed:With regard to the assumptions concerning the specimens’ boundary conditions and the technical conditions of the tests performed for the calculation, the tested elements were positioned between a rigid support at the bottom and a hinge joint suspended to the machine travers at the top. Data cables were placed through an opening in the steel plate for load transfer at the rigid support. In order to improve the condition of the contact of the specimen with the rigid materials, two PTFE intermediate layers with thicknesses of 2 mm were installed between the steel surfaces of the testing device and the specimen edges. In The top and bottom edges of all the hollow cylinders were cut and ground to ensure a constant length of every specimen and perpendicularity between the edge surface of the cylinders and their longitudinal axis.Electrical resistance wire strain gauges, each with a length of 75 mm, were glued on to the concrete surface and connected with the use of data cables to the Spider 8 HBM data capturing system. At the time of the tests, all the cylinders  were prepared with six strain gauges glued on the internal surface at its mid-length—three gauges in the circumferential direction and three gauges in the longitudinal direction. In the same area of the concrete element in the reference specimen (R3), a set of strain gauges was also installed on the external surface, using the same layout as before. Moreover, at a distance of 75 mm from the upper edge of this specimen, there were two more strain gauges installed in the circumferential direction. The selected position was chosen in relation to the prestressed specimens (R1 and R2) as it was halfway between the specimen edge and the first turn of the wires. In order to limit the complications of connecting the strain gauges to the capturing system, the additional group of gauges was installed only on the external surface of the hollow cylinder.In the hollow cylinders termed R1 and R2, beside the strain gauges at their mid-length there were also additional gauges installed at a distance of 75 mm from the upper edge of the specimen, in both circumferential and longitudinal directions. Additionally, two more gauges were also installed in the bottom part of the hollow cylinders, at the halfway point between the wire windings and the edge.The preparation of the specimens for the tests was as follows:The reference specimen termed R3 was positioned as the first in the testing machine and compressive force was applied. The maximal value of the force applied to the sample was 293 kN at a vertical displacement of the loading surface of 3 mm. In the next step, R1 and R2 were tested. The displacement increase rate was kept constant for all the specimens. In In The external surface of the non-prestressed reference specimen R3 was the most damaged—the number of cracks in the horizontal and in vertical directions is higher than on the other cylinders. The majority of the cracks were concentrated in the area close to the bottom edge of the cylinder (some were more than 100 mm from the edge) and individual vertical cracks were up to 300 mm from the edge The cracked area in the two other tested samples (R1 and R2) was visibly smaller. Vertical cracks generated at the bottom edge of the hollow cylinders did not reach the lowest wire turn. Nevertheless, other instances of splitting and spalling in the bottom part of the specimens were more severe, as shown in maxP = 293 kN at t = 3485 s;for the reference cylinder R3—oP = 640 N)—maxP = 282 kN at t = 4750 s;for the hollow cylinder R1 (prestressed with wire tensioned to oP = 1280 N)—maxP = 224 kN at t = 5756 s.for the hollow cylinder R2 . During the final phase of the test, at t = 66,878 s, the mean vertical strain in the concrete measured in the bottom part of the tested hollow cylinder was 1346 μm/m. Significant deformations close to the bottom edge were observed together with vertical cracks (The final compressive strain in the concrete in specimen R2 measured in two directions on the internal surface at the mid-length point and on the external surface at the upper part of the hollow cylinder remained similar, with values of 724 μm/m and 766 μm/m see c. The stl cracks b and thet = 2293 s and with a loading force of P = 206 kN, the circumferential tensile strain of the concrete increased significantly. A maximal tensile strain value of −772 μm/m was captured at the end of the measurement period, just after the moment when a crack generated at the bottom edge developed rapidly and reached halfway along the hollow cylinder, destroying one of the strain gauges (t = 3000 s. With further loading increases, deformations in the R1 hollow cylinder depend on prestressing action. Circumferential strain increases to a maximal value of −175 μm/m at 4741 s, which corresponds to a loading force of 282 kN. Further increases in the loading to its maximal value P = 282 kN did not substantially affect the strain, the measured value of which at the final step at the time 5226 s from loading was −174 μm/m.In n gauges a. The vat = 3295 s at a loading force of P = 133 kN. With further load increases to the maximal level of P = 224 kN at the time t = 5135 s, the mean circumferential strain was −40 μm/m. The final circumferential strain of the concrete after 6226 s was −54 μm/m.Observing the behavior of the hollow cylinder R2, it was concluded that the maximal tensile strain in the circumferential direction at the level of −97.8 μm/m was observed after a loading time of P = 105 kN was reached on the internal surface of the concrete wall, circumferential compressive strain was measured (see P = 150 kN, the mean values of the circumferential concrete strain measured on both surfaces were equal to each other. During the final step of the measurements, the maximal circumferential tensile strain on the internal and external surfaces of the R3 cylinder were −772 μm/m and −343 μm/m, respectively. At the upper measurement location of the R3 specimen, the circumferential strain of the concrete was tensile during the whole testing time and its value reached −533 μm/m at its final phase of t = 3784 s.The behavior of the R3 cylinder was different as in the first 1470 s of the test up until the loading force of ured see a, while ured see a. SubseqA comparison of the final values of the circumferential tensile strain shows that for the R1 hollow cylinder after 5526 s of tests, the value measured in its upper part was −286 μm/m; and for the R2 hollow cylinder after 6878 s, the value was −182 μm/m. For both specimens, the measured values in the upper area were higher than the comparable strain values on the internal surface of the mid-length point of the hollow cylinder. Measurements taken in the bottom part of the R1 hollow cylinder for the final circumferential strain on the external surface after 4320 s of the test were −1140 μm/m. Comparable strain values for the R2 specimen were at −1120 μm/m after the full test time of 6878 s.The observed results of the concrete strain clearly reveal the influence of the circumferential prestressing of hollow concrete cylinders on their behavior. In the upper part, the mean tensile strain of concrete in the circumferential direction measured on the external surface of the R3, R1 and R2 specimens at the final phase of the test were −533 μm/m, −286 μm/m and −182 μm/m, respectively. The mean strain in the same direction in the lower part of specimens R1 and R2 were at the same level, with values of −1140 μm/m after a loading time of 4320 s and −1120 μm/m after a loading time of 6878 s. During this phase of the test, concrete spalling together with buckling of the bottom part of the hollow cylinder in the outside direction was observed.Changes in the mean values of the circumferential strain of the concrete measured on the internal surface of the hollow cylinders at the mid-length as a function of the loading time shown in Variations in the mean concrete strain in the circumferential direction measured on the internal wall surface at its mid-length point in relation to the compression loading for specimens R1, R2 and R3 are presented in Higher prestressing applied to the R2 specimen resulted in the limited development of the circumferential strain of concrete in the section above the prestressed area. The transversal confinement of the hollow cylinder lengthened the loading diagram and allowed higher displacement of the upper loading surface in the testing machine, assisted by larger deformation of the concrete at the bottom edge of the specimen. This deformation of concrete was critical and limited the maximal load applied to the specimen during the test.Diagrams of the mean strain of concrete in the vertical direction in relation to the vertical compressive force measured halfway along the length of the internal surfaces of the R1, R2 and R3 hollow cylinder specimens are shown in Finally, it should be noted that the responses of specimens R1 and R2 were different (see The application of SMA wires for the prestressing of the hollow concrete cylinders provided wider information on their effectiveness in comparison to plain cylinders prestressed with a similar method. This may be used in the future, followed by further tests to establish design provisions for this method of strengthening structural members, mainly columns.The application of SMA wires wound directly over concrete members in order to strengthen them cause the concrete to heat, which may lead to tensile stress in line with the concrete’s tensile strength.On the basis of the research data reported in the bibliography, it may be stated that the prestressing of hollow cylinders with SMA wires that need longer heating for the activation of the SM effect can cause the destruction of the prestressed element.For the reason described above it is advised that prestressing programs using this technique should hinder concrete heating by allowing the wire to remain in physical contact. Lower-diameter Ni-Ti wires and larger spacing of the wire windings are therefore preferred. The analysis of the influence of the wire diameter and the spacing of the windings on the temperature increase of the concrete is a potential direction of further research in this area.The tests on the hollow concrete cylinders performed one year after the application of prestressing by means of Ni-Ti wire confirmed that the residual stress in the wire remained present during the whole time period. It may therefore be concluded that the strengthening of a concrete element designed with use of Ni-Ti SMA material is effective for a longer time. At the same time, it should be underlined that strengthening should be applied along the entire length of the member without allowing any uncompressed sections.The following conclusions can be drawn from the study:


Pseudomonas aeruginosa (Pa) is a major bacterial pathogen responsible for chronic lung infections in cystic fibrosis patients. Recent work has implicated Pf bacteriophages, nonlytic filamentous viruses produced by Pa, in the chronicity and severity of Pa infections. Pf phages act as structural elements in Pa biofilms and sequester aerosolized antibiotics, thereby contributing to antibiotic tolerance. Consistent with a selective advantage in this setting, the prevalence of Pf-positive (Pf+) bacteria increases over time in these patients. However, the production of Pf phages comes at a metabolic cost to bacteria, such that Pf+ strains grow more slowly than Pf-negative (Pf−) strains in vitro. Here, we use a mathematical model to investigate how these competing pressures might influence the relative abundance of Pf+ versus Pf− strains in different settings. Our model suggests that Pf+ strains of Pa cannot outcompete Pf− strains if the benefits of phage production falls onto both Pf+ and Pf− strains for a majority of parameter combinations. Further, phage production leads to a net positive gain in fitness only at antibiotic concentrations slightly above the MIC  but which are not lethal for Pf+ strains. As a result, our model suggests that frequent administration of intermediate doses of antibiotics with low decay rates and high killing rates favors Pf+ over Pf− strains. These models inform our understanding of the ecology of Pf phages and suggest potential treatment strategies for Pf+ Pa infections.IMPORTANCE Filamentous phages are a frontier in bacterial pathogenesis, but the impact of these phages on bacterial fitness is unclear. In particular, Pf phages produced by Pa promote antibiotic tolerance but are metabolically expensive to produce, suggesting that competing pressures may influence the prevalence of Pf+ versus Pf− strains of Pa in different settings. Our results identify conditions likely to favor Pf+ strains and thus antibiotic tolerance. This study contributes to a better understanding of the unique ecology of filamentous phages in both environmental and clinical settings and may facilitate improved treatment strategies for combating antibiotic tolerance. Pseudomonas aeruginosa (Pa) infections, particularly those that are antibiotic resistant or antibiotic tolerant, are responsible for growing health care expenses and considerable mortality (Pa infections are particularly problematic in cystic fibrosis (CF), an inherited disease associated with thick, tenacious airway secretions  strains, while over 40% of adults and 100% of 10 prelung transplant patients have Pf+ strains . This suggests that infection of Pf-negative (Pf−) strains from a Pf+ strain is rare. Together these data point to selective advantages for Pf+ Pa strains in the setting of chronic infection. However, the production of Pf phages comes at a steep metabolic cost to the Pa strains that produce them. Pf phages are abundantly expressed in Pf+ Pa biofilms ]—due to the extra energetic cost θ of phage production for Pf+ strains.The metabolic cost of phage production θ leads to a reduction in Pf+ strains per-capita growth rate and, ultimately, in a lower long-term equilibrium for Pf+ strains relative to Pf− strains . When inCF lungs . Pf+ and2 CFU/ml) of Pf− bacteria is introduced and directly competes with an established Pf+ bacterial colony, assuming θ> 0, the density of the Pf− strain increases over the course of multiple days until it reaches a density close to that of the Pf+ strain .When competing directly for space and resources, Pf+ and Pf− bacterial strains experience the same phage and antibiotic concentrations in their environment. In this section, we assess the antibiotic effect on both strains living in direct competition, and we explore two antibiotic scenarios, namely: when the antibiotic effect is not dependent on the replication rate of the target bacteria εr = 0, , and wher = 0), the population growth rate of Pf− strains remains higher than the growth rate of Pf+ strains across the range of antibiotic concentrations (When the antibiotic effect is not dependent on replication rate (εtrations . In thisr = 1). In such a case, the benefit of phages’ protective biofilm is still equally shared between Pf+ and Pf− strains. However, the energetic cost θ of phage production  The Pf+ strain is more likely to outcompete the Pf− strain at intermediate antibiotic concentrations (20× MIC). In the case of nonoverlapping bacterial strains (indirect competition), phages are absent in Pf− strain environments, and thus, there is no change in the effective antibiotic concentration experienced by Pf− strains. In Pf+ bacterial infection sites, in contrast, phages can sequester some amount of the antibiotics, essentially lowering the local concentration of antibiotics acting on Pf+ bacteria.2 = 5.11; P = 0.024; confidence interval [CI], 0.047 ± 0.041). Both Pf+ and Pf− strains go extinct in the remaining 42% of runs. At this intermediate antibiotic concentration, there is more variation between runs, with some parameter sets leading to rapid extinction, while others lead to slow declines over multiple days or oscillate under the antibiotic-free density .At low antibiotic concentrations (doses of 0.1× MIC), the Pf− strain maintains a higher density than the Pf+ strain over the last day of treatment, for 84% of parameter sets . Both st density  and F. F density  to I. At(ii) The fitness of the Pf+ strain relative to the Pf− strain is determined by different parameters at different antibiotic concentrations. We performed a partial rank coefficient analysis to determine which parameters had the strongest relationships with Pf+ density, Pf− density, and the ratio of Pf+ to Pf−  or set to 0.1× MIC (B), 5× MIC (C), or 20× MIC (D). Download Copyright © 2021 Pourtois et al.2021Pourtois et al.https://creativecommons.org/licenses/by/4.0/Creative Commons Attribution 4.0 International license.This content is distributed under the terms of the 10.1128/mSystems.00193-21.2TABLE S1P values for the hypothesis that there is no partial correlation between each parameter and Pf+ density, Pf− density, and the ratio between Pf+ and Pf− density, falling under four different possible ranges: below 0.01 (***), between 0.01 and 0.05 (**), between 0.05 and 0.1 (*) and above 0.1 (-). Download Table S1, DOCX file, 0.02 MB.Copyright © 2021 Pourtois et al.2021Pourtois et al.https://creativecommons.org/licenses/by/4.0/Creative Commons Attribution 4.0 International license.This content is distributed under the terms of the B were negatively correlated with the ratio of Pf+ to Pf− density  for the intercept and slope of each linear model. Each model includes an intercept, a parameter, antibiotic concentration, a quadratic term for antibiotic concentration, and an interaction between the parameter and antibiotic concentration. Download Copyright © 2021 Pourtois et al.2021Pourtois et al.https://creativecommons.org/licenses/by/4.0/Creative Commons Attribution 4.0 International license.This content is distributed under the terms of the In addition to metabolic cost and antibiotic sequestration, phage production and antibiotic decay were both positively correlated with Pf+ relative fitness when 5× MIC was used . Up to 2(iii) High initial bacterial density is necessary for phages to provide nonnegligible benefits. Pf+ relative fitness slowly increases as the initial bacterial concentration (before the start of the antibiotic treatment) increases, with a sharp increase at densities close to the carrying capacity of 108 CFU/ml  sequester 80% of antibiotics at a concentration of 1× MIC, but this measure drops sharply as phage density decreases  are critical for Pf+ strains to significantly outcompete Pf− strains. These concentrations correspond to the regimen at which antibiotic sequestration by phages lowers the effective antibiotic concentration—thus lowering the death rate of Pf+ bacteria—enough to compensate for the higher energetic cost of phage production. Under these circumstances, Pf+ bacterial density exceed that of Pf− bacteria. Below these antibiotic concentrations, the marginal benefit of a reduction in antibiotic effective concentration accrued thanks to the protective effect of the phage biofilm is unable to compensate for the cost of phage production, and thus Pf+ density remains lower than that of Pf− strain. At high antibiotic doses, the ability of phages to sequester antibiotic is overwhelmed, and infections with both Pf+ and Pf− are eventually cleared. This is consistent with the clinical evidence that shows Pf+ strain infections are associated with chronic infections  that havPa survival. For example, Pf phages may impact the rheological properties of the infection environment, the adhesivity of the biofilm, the transport of nutrients and antibiotics, as well as any immune response. In turn, these properties can affect the physiological state of bacteria, and phage production  types of interaction between Pf+ and Pf− strains: (i) direct competition in a mixed population of Pf− and Pf+ bacterial strains coexisting in the same location  and (ii) no direct interaction with Pf− and Pf+ strains coexisting in the same patient but established in different  locations, e.g., spatially separated infection sites in the lungs —hereafteIn direct competition, both Pf− and Pf+ strains must compete for the same nutrients and space. As filamentous phages are generally highly specific , 65, we In indirect competition, there is no interstrain competition for nutrients and space, and we assumed that, under antibiotic treatment, the Pf− strain does not benefit from the protective effect of phages produced elsewhere by the Pf+ strain.(i) Bacterial growth. In the absence of antibiotics, bacteria iB (in CFU/milliliter) from each strain i = Pf+ or Pf−, replicate at a density-dependent rate Φ(Btot) and die at the per-capita rate δB:Btot) as a decreasing logistic function of the total bacterial concentration in the local biofilm Btot:rmax is the absolute maximum replication rate, K is the bacterial density at which the per-capita reproductive rate Φ(Btot) is equal to zero, and θ is the reduction in the per‐capita replication rate caused by the metabolic cost of phage production in Pf+ bacterial strains only . When both strains are competing directly for space and resources, the total bacterial concentration Btot is the sum of their concentrations, namely: Btot = B+ + B−. When competing indirectly, Btot is set equal to the local concentration of bacteria, i.e., either the concentration of the Pf+ strain or the Pf− strain.We choose to model bacterial growth directly with a logistic equation rather than a Gompertz equation or a mechanistic approach with nutrients, as it only requires estimating the maximum growth rate and carrying capacity. This model and its variants are used regularly to describe bacterial growth , 66, 67.(ii) Phage production. Filamentous phages V (PFU/milliliter) are produced by Pf+ bacteria at a constant rate λ and decay at a rate δV:V were numerically optimized (see “Parametrization”) to obtain viral densities an order of magnitude larger than bacterial densities, consistent with literature reports of Pa infections (Both λ and δfections .(iii) Antibiotic treatment. We consider a range of antibiotic regimens, with different antibiotic concentrations per dose, administered twice a day. We assume each dose of antibiotics leads to an instantaneous peak Amax in the sputum. All concentrations discussed in this work represent concentrations in the sputum, which can differ in nontrivial ways from the original concentration administered orally. After administration, antibiotics are removed from the system via both degradation of the drug and clearance due to natural flow through the body. Here, we assumed that their concentration follows first order kinetics  is the per-capita reduction in bacterial growth rate or, equivalently, the increase in mortality rate as a consequence of antibiotics at concentration A, bacterial density B, and phage density V—hereafter referred to as the antibiotic killing rate. Ψ is here described by a Hill function  that levels off to A50(B) is the antibiotic concentration leading to half H, the Hill parameter, is proportional to the steepness of the function at Aeff = A50 ψ and bacterial initial density on Pf+ relative fitness. We calculated the difference between the proportion of runs dominated by Pf+ bacteria and the proportion dominated by Pf− bacteria for a range of each parameter and antibiotic concentration. We then fit a linear model with an interaction term between parameter and antibiotic concentration and a quadratic term for antibiotic concentration. Results are reported in To assess the sensitivity of the outcome to model parameters, we performed a partial rank correlation coefficient analysis (PRCC) with Latin hypercube sampling (LHS) after log transformation of model parameters over the ranges reported in https://github.com/jpourtois/phage-antibiotics.All code is available on GitHub at


Respiratory distress syndrome (RDS) is a condition caused by a deficiency in pulmonary surfactant. Many interventions, including pulmonary surfactant, non-invasive respiratory support, and other supportive treatments have been used to prevent RDS. However, recent studies have focused on the continuous positive airway pressure as a significant potential agent for preventing RDS. However, its safety and effectiveness are yet to be assessed. To this end, the current study aims to perform to explore the safety and effectiveness of continuous positive airways in treating neonates with RDS.We will conduct comprehensive literature searches on MEDLINE, Cochrane Library, EMBASE, Chinese National Knowledge Infrastructure, and Chinese BioMedical Literature from their inception to April 2021. The search aims to identify all the randomized controlled studies on continuous positive airway pressure in treating neonates with RDS. In addition, we aim to search the gray literature to establish any available potential studies. We will use 2 independent authors to determine study eligibility, extract data using the structured pro-forma table, analyze data, and utilize suitable tools in assessing the risk of bias in the selected studies. Accordingly, we will conduct all statistical analyses using RevMan 5.3 software.The current study aims to provide high-quality synthesis of existing evidence concerning the continuous positive airway pressure to treat neonates suffering from RDS.Our findings seek to provide evidence to establish whether continuous positive airway pressure can ascertain safety and effectiveness for neonates with RDS.The study will require ethical approval.https://osf.io/7nj8s/).May 20, 2021.osf.io/7nj8s. ( Accordingly, the physiologic role of surfactant entails the capacity to lower surface tension. It also has the capacity of rapidly absorbing, spreading, and reforming a monolayer in the dynamic conditions related to the respiratory cycle. At the same time, the administration of exogenous surfactant can moderate mortality and risk of air leak, especially since it is regarded as the basis of therapy for preterm infants with RDS.,6Respiratory distress syndrome (RDS) is instigated by a deficiency or dysfunction of pulmonary surfactant. The surfactant lines the alveolar surface, preventing atelectasis at end-expiration. Accordingly, a pulmonary surfactant is considered an active agent capable of keeping the pulmonary alveoli open as well as facilitating the entry of air to the lungs to enhance oxygenation in neonates.,8 Drawing from existing studies, it is evident that application of nasal continuous positive airway pressure from birth is considered effective as intubation and ventilation for newborns, especially those that are less than 30 weeks growth.–11 However, applying continuous positive airway pressure from the beginning in an unselected population of preterm newborns can cause the risk of under-treating of those infants RDS and for those whom continuous positive airway pressure might fail to afford sufficient respiratory support. At the same time, the effectiveness of continuous positive airway pressure is well-known extensively acceptable among clinicians; however, whether patients benefit more or less from continuous positive airway pressure is less clear. Still, continuous positive airway pressure has been approved to be used for neonates with RDS, meaning that it is considered crucial to understanding the efficacy and safety of continuous positive airway pressure to treat neonates with RDS. Therefore, our study aims to evaluate the safety and effectiveness of continuous positive airway pressure in treating neonates with RDS.In the recent past, non-invasive respiratory support has become more and more prevalent for respiratory dysfunction management, especially among preterm newborns.2The present study seeks to explore the safety and effectiveness of continuous positive airway pressure in treating neonates suffering from RDS.33.1https://osf.io/) under number 10.17605/OSF.IO/7NJ8S. It follows the guidelines put forward by Preferred Reporting Items for Systematic Review and Meta-Analysis Protocol statement.The present protocol was registered in OSF , particularly those considered to be at risk of developing RDS of those with an already existing clinical diagnosis of RDS.3.2.3We will include continuous positive airway pressure treatment with any pressure level. We will also consider those delivered by any type of device, for any given duration, and presented in any mode, compared to no-continuous positive airway pressure or sham continuous positive airway pressure. We will also include studies that applied continuous positive airway pressure at any time after the patient presentation.3.2.4The major outcomes for the present study are the number of neonates needing mechanical ventilation as well as recovery time. However, the minor outcomes will include changes in the respiratory rate, change in the saturation of the arterial oxygen, emergency department stay duration, hospital stay duration, the requirement for admission in the intensive care unit, and other adverse incidents.3.3We will carry out an in-depth literature search on MEDLINE, Cochrane Library, EMBASE, Chinese National Knowledge Infrastructure, and Chinese BioMedical Literature from their inception to April 2021, to identify all the randomized controlled studies on continuous positive airway pressure in treating neonates suffering from RDS. We will employ the following search terms singly or as combinations: “respiratory distress syndrome,” “continuous positive airway pressure,” and “randomized controlled trial.” Besides, we intend to search the gray literature to identify any potential studies.3.43.4.1We intend to include all randomized and quasi-randomized controlled trials that will meet the established selection criteria prescribed in the previous section. Accordingly, we will use 2 independent authors to review the search results and independently select studies that meet the inclusion criteria. Also, we will resolve any disagreements between the authors through discussion and by involving a third author. Figure 3.4.2We will use 2 independent authors to also extract, assess, and code all data obtained from the search, for every study, through the use of a pilot data extraction designed purposefully for our study. Also, we plan to utilize a third author to resolve any disagreements between the 2 independent authors. Also, the authors will independently enter data into RevMan 5.3, with another author checking the entries.3.4.3 In case of disagreements, we intend to resolve them through discussion with a third author.Furthermore, we plan to employ 2 independent authors to evaluate the risk of bias of all included trials by utilizing the Cochrane “Risk of bias” tool.3.4.4We will analyze continuous outcome data using the mean differences (MDs) or standardized MDs. Also, we will analyze dichotomous outcome data by utilizing the relative risk. In particular, we intend to report the 95% confidence intervals on all estimations.3.4.5We will rely upon evaluators to verify key study features and attain missing numerical outcome data where pertinent.3.4.6I2 statistic in measuring the heterogeneity of the studies in every analysis. Where we find significant heterogeneity , we intend to use the random-effects models; otherwise, we will use the fixed-effects model.,15Our study will also rely on utilizing the 3.4.7Our study will also consider creating and examining a funnel plot to examine possible small studies as well as publication biases where more than 10 studies are included.3.4.8Furthermore, our study will perform sensitivity analysis to investigate our findings’ reliability and robustness.4The present study aimed at assessing the safety and effectiveness of continuous positive airway pressure in treating neonates with RDS. Recently, the randomized controlled trials of continuous positive airway pressure for the treatment of neonates suffering from RDS have increased gradually. While many published studies seem to suggest that the application of continuous positive airway pressure has a significant position to treat newborns suffering from RDS, its safety and effectivity in treating neonates with RDS is still indecisive. Therefore, our study seeks to explore this topic, assessing the safety and effectivity of continuous positive airway pressure in treating neonates with RDS. We are confident that our study presents a practical guide for practitioners to make valid decisions when treating neonates with RDS. Our study will also be critical for health policymakers.Conceptualization: Junyi Cao, Zuowu Chen.Data curation: Junyi Cao, Zuowu Chen, Jinbing You.Formal analysis: Junyi Cao, Jinbing You.Funding acquisition: Jiangjiang Wang, Qiong-Yao Tang.Investigation: Jiangjiang Wang.Methodology: Junyi Cao, Jinbing You, Jiangjiang Wang.Project administration: Jiangjiang Wang, Qiong-Yao Tang.Resources: Junyi Cao, Jiangjiang Wang, Qiong-Yao Tang.Software: Junyi Cao, Zuowu Chen.Supervision: Jinbing You, Jiangjiang Wang.Validation: Zuowu Chen.Visualization: Zuowu Chen.Writing – original draft: Junyi Cao, Zuowu Chen.Writing – review & editing: Jiangjiang Wang, Qiong-Yao Tang.


Multiaxis three-dimensional (3D) continuous polymeric carbon fiber/cementitious concretes were introduced. Their angular (off-axis) flexural properties were experimentally studied. It was found that the placement of the continuous carbon fibers and their in-plane angular orientations in the pristine concrete noticeably influenced the angular flexural strength and the energy absorption behavior of the multiaxis 3D concrete composite. The off-axis flexural strength of the uniaxial (C-1D-(0°)), biaxial (C-2D-(0°), and C-2D-(90°)), and multiaxial (C-4D-(0°), C-4D-(+45°) and C-4D-(−45°)) concrete composites were outstandingly higher (from 36.84 to 272.43%) than the neat concrete. Their energy absorption capacities were superior compared to the neat concrete. Fractured four directional polymeric carbon fiber/cementitious matrix concretes limited brittle matrix failure and a broom-like fracture phenomenon on the filament bundles, filament-matrix debonding and splitting, and minor filament entanglement. Multiaxis 3D polymeric carbon fiber concrete, especially the C-4D structure, controlled the crack phenomena and was considered a damage-tolerant material compared to the neat concrete. Textile concrete composite (TCC) consists of two-dimensional 2D) biaxial fabric or three-dimensional 3D) preform with multiple continuous fibers as the reinforcement materials and fine-grained cement as the matrix material. TCC can be considered noncorrosive, with better fracture toughness probably due to crack control via unusual stress transfer mechanism and lightweight compared to the steel concretes. Research on the TCC especially two-dimensional (2D) fabric reinforced concretes, was reported by several well-known researchers ,2,3,4,5.D biaxialD preformTextile technologies were used to make various 2D and 3D structures to apply in concrete end-uses ,8. Non-iIt has been observed that concrete beams with bars made of carbon fiber-reinforced polymer (FRP) exhibited worse resistance to cracking and faster crack propagation during the entire flexural loading process than steel bars. Flexural shear failure was observed due to the cracking patterns in FRP-based concrete bars . HoweverGeopolymer cement, which was alkaline-based, including fly ash or slag from iron and metal, and carbon fiber bars concrete demonstrated low crack width than the ordinary Portland cement-based concrete . BiaxialCarbon and glass fabric reinforced mortar with micro steel fiber was reported to enhance the energy dissipation capability of concrete panels, but their flexural properties were not improved . The addThe 3D glass woven preform was interlaced to form a thick substrate incorporating polyurethane/epoxy and cementitious matrix for making textile concrete composite . The fleIn the case of textile reinforced mortar (TRM) jackets, the U-wrapping strengthening configuration is much more effective than the side-bonding configuration, and increasing the number of layers significantly increases the effectiveness of the TRM jacket . AnotherThe fracture toughness of the cementitious matrix was increased by debonding fiber-matrix interfaces, fiber entanglement, pulling out, and stick-slip phenomena. Thus, fibers in the concrete structure administered the micro/macro crack propagation ,65. The The artificial neural network (ANN) method was used to model the flexural strength of TCC where concrete with carbon fiber fabric and short fibers was used to improve the initial crack and post crack load capacity . Carbon Numerous research studies have been carried out on concrete with added staple fiber or fabric substrate. However, limited research on the mechanical properties of the 3D concrete structure was examined considering a dry, continuous, filamentous TOW-based layered architecture with large meshes. Thus, the research aimed to develop multiaxially oriented high-strength and high modulus continuous carbon fiber-based concrete composite and to search experimentally the principal and angular (off-axis) flexural properties of these concrete composites.3, and 6 μm, respectively.Polyacrylonitrile (PAN) based carbon fiber  were employed to design the multiaxis three-dimensional (3D) preforms. The properties of the carbon fibers are exhibited in 2 based silica, according to AS 114131, fine aggregate 100–400 μm, density 2.65 g/cm3, melting temperature 1750 °C, Pomza export, Manisa, Turkey), superplasticizer  to enhance the workability and rheology, and water (city water) was used to produce the multiaxis three-dimensional (3D) continuous carbon fiber reinforced concretes.A cementitious matrix including cement , siliceous fine sand  preform structures were chosen for the study due to eliminating interlacement-based crimp exchange and crimp extension (or displacement), as in the case of biaxial fabric . This prIn the flat winding/molding rig, the fibers, which were plied under low twists for increasing the volume fraction and prevent spreading the filaments, were multiaxially, biaxially, uniaxially oriented without any interlacements. As indicated in 4, [90°/0°]2, and [90°/±45°/0°]1, respectively.All the multiaxis 3D carbon concrete structures  was principally composed of four layers with different directions such as uniaxial (0°), biaxial (0°/90°), and multiaxial (90°/±45°/0°) direction. Some of the multiaxis 3D carbon concrete composites are shown in 3, 34.51, wt.%) and siliceous fine sand  were weighed and then mixed at 100 rpm for a minute  and superplasticizer  were added with the dry mixture and mixed for one minute at 150 rpm [The PAN carbon filamentary TOWs preforms were consolidated via hand lay-up to make multiaxis 3D concrete composites. The multiaxis 3D concrete composite processing steps are shown schematically in a minute . After t 150 rpm . Then a ® software with a 100 kN loading cell. The cross-head speed was 0.9 mm/min. The flexural tests in the principal and ±bias directions were repeated 5 times and 3 times, respectively. The flexural test was done at the standard laboratory atmosphere condition (temperature of 23 ± 2 °C and relative humidity of 50 ± 10%). The ASTM C642-13 standard was followed to have the sample’s density, absorption, and voids [psσ) [sI).b is the sample width (mm), d is the sample thickness (mm), L is the support span (mm); P is the maximum applied load (N); psσ is the flexural strength (MPa); fcP is the first crack load (N), fcδ is the first crack displacement (mm) and sI is the initial stiffness (N/mm).A flexural test using four-point loading was conducted to determine the flexural properties of the multiaxis 3D carbon concrete composite panels. ASTM C78/C78M-18  and ASTMnd voids . In addind voids . Equatiods [σps) . AdditioA scanning electron microscope  was used to analyze the morphology of the samples before and after the flexural test. The flexural samples’ macro-scale images were identified using a digital camera . A steel ruler was employed to measure the crack lengths in the fractured flexural samples. A digital caliper  were utilized to measure the crack widths in the fractured flexural samples. A digital ruler was employed to measure the crack distance on the sample under four-point flexural load . Off-axi® software. Therefore, statistical computations on the raw data, including flexural first crack load and displacement, flexural maximum load and displacement, flexural first crack strength, and flexural maximum strength, as well as flexural stiffness, were carried out and analyzed as presented in After the flexural test data for the pristine (CC) and carbon concrete (C-1D-(0°), C-1D-(90°), C-1D-(+45°), C-2D-(0°), C-2D-(90°), C-2D-(+45°), C-4D-(0°), C-4D-(90°), C-4D-(+45°) and C-4D-(−45°)) obtained from the Shimadzu AG-X 100 (JP) testing machine, they were transferred to the Microsoft (MS) Excel spreadsheet 2013 via TrapeziumIn In In As depicted in a general characterization of the load-displacement curves of C-1D, C-2D, and C-4D PAN carbon TOW concretes , the fibThe average maximum flexural load, first crack load, maximum load displacement, and first crack displacement for the developed multiaxis 3D carbon concrete composites are shown in For the flexural maximum load in principal directions (0° and 90°), C-1D-(0°) was 41.62% higher than C-2D-(0°) and was 63.94% higher than for C-4D-(0°), and C2D-(0°) was 15.76% higher than C-4D-(0°). However, C-4D-(90°) was 88.82% lower than C-2D-(90°) and was 14.42% higher than C-1D-(90°), and C2-D-(90°) was 2.16 fold greater than C-1D-(90°). For the flexural maximum load in the bias directions (+45° and −45°), C-4D-(+45°) was 48.79% greater than C-2D-(+45°) and was 117.13% greater than C-1D-(+45°), and the C-2D-(+45°) was 45.93% higher than C-1D-(+45°). In addition, C-4D-(+45°) was 30.99% lower than C-4D-(−45°). It was found that the flexural load carrying performance on various developed carbon fiber concrete composites was proportional to the fiber orientation and amount of fiber in a particular direction. In addition, the flexural load-displacement of the C-1D-(0°), C-2D-(0° and 90° and +45°) and C-4D-(0° and ±45°) concrete composites were increased from 36.84 to 272.43% compared to the pristine concrete. The flexural performance of the concretes was influenced by the placement and orientation of the continuous PAN carbon fiber. In the first crack load , the firAs depicted in Average flexural maximum strength, first crack strength, and initial stiffness for the developed multiaxis 3D carbon concrete composites are shown in For the flexural maximum strength in the principal directions (0° and 90°), the C-1D-(0°) was 41.65% higher than C-2D-(0°) and was 63.95% higher than C-4D-(0°), and the C-2D-(0°) concrete composite was 15.74% higher than C-4D-(0°). However, the C-4D-(90°) concrete was 88.82% lower than C-2D-(90°) and was 14.49% higher than C-1D-(90°), and the C-2D-(90°) concrete was 2.16 fold greater than C-1D-(90°). For the flexural maximum strength in the bias direction (+45° and −45°), the C-4D-(+45°) was 44.28% greater than C-2D-(+45°) and was 2.17 fold greater than C-1D-(+45°), and C-2D-(+45°) was 45.78% higher than C-1D-(+45°). In addition, the C-4D-(+45°) composite was 31.08% greater than the C-4D-(−45°). It was found that the flexural strength performance of various developed carbon fiber concrete composites was proportional to the fiber orientation and amount of fiber in a particular direction, depending upon the structural fiber architecture. In addition, the flexural strength of the C-1D-(0°), C-2D-(0° and 90° and +45°) and C-4D-(0° and ±45°) concrete composites were 272.54 to 36.83% higher than that of the neat concrete, respectively. It was discovered that the placement and orientation of the continuous PAN-based carbon fiber in the neat concrete influenced the flexural strength performance of the concrete. As exhibited in As shown in ffcE is the flexural first crack energy corresponding to the load-displacement in area C0C1C2; fmlE is the flexural maximum load energy corresponding to the load-displacement in area C0C1C3C4; fshE is the flexural strain hardening energy corresponding to the load-displacement in area C2C1C3C4; fssE is the flexural strain-softening energy corresponding to the load-displacement in area C4C3C5C6, and ftE is the flexural total energy corresponding to the load-displacement in area C0C1C3C5C6.As illustrated in As exhibited in For the PAN-based carbon fiber concretes , the flexural strain hardening energy ranged from 3.88 to 11.01 J. Considering the flexural strain hardening energy in the principal directions (0 and 90°), the C-1D-(0°) concrete was 82.59% higher than C-2D-(0°) and was 2.84 fold higher than C-4D-(0°), and the C-2D-(0°) composite was 55.41% higher than C-4D-(0°). However, the C-2D-(90°) concrete was extraordinarily higher compared to C-4D-(90°), and the C-1D-(90°) had no strain hardening energy stage. For the flexural strain hardening energy in the bias directions (+45° and −45°), the C-4D-(+45°) was 2.28% higher than C-4D-(−45°). However, the C-2D-(+45°) and C-1D-(+45°) composites had no strain hardening energy stages. It was found that the flexural strain hardening energy absorption performance on various developed carbon fiber concrete composites was proportional to the fiber orientation and amount of fiber in a particular direction, depending upon the structural fiber architecture. In addition, the C-1D-(0°), C-2D-(0° and 90°), and C-4D-(0° and ±45°) concrete composites were extraordinarily higher compared to the pristine concrete. The Fourier transform infra-red (FTIR) spectra of the CC (pristine concrete), PAN-based carbon fiber, C-1D, C-2D, and C-4D carbon fiber reinforced concretes are shown in −1 was assigned to the C–O stretching vibration. Concerning the other parts of the spectrum, the band at 1083 cm−1 may be attributed to Si-O asymmetric stretching vibrations [−1 can be attributed to the -C-C-, -C-H- groups [−1 was the absorption peak of the Si-OH and –OH groups. In the FTIR spectrum of the PAN carbon fiber (C), as illustrated in −1 were the deformation vibrations of the -CH2-CH-  groups, whereas the peak at 2336 cm−1 was ascribed to the deformation vibrations of the functional group of −C≡N (nitryl) [−1 and 1356 cm−1 were attributed to the stretching vibrations of the -C=N- group. In addition, the peaks observed between 1194 and 885 cm−1 corresponded to -C-C- and -C-N- bonds [−1 bands may be attributed to the deformation vibrations of H2O, polycarboxylate, -Si-OH, and -Ca-OH groups. In the 3000–2850 cm−1 band spacing, the deformation vibrations of -CH2-CH-  groups were observed, whereas the signal between 2300 and 2200 cm−1 corresponded to the −C≡N (nitryl) group. The signals between 1600 and 1300 cm−1 were attributed to stretching vibrations of the -C=C-, -C=O, and -C=N groups. The signal observed between 1100 and 1000 cm−1 was assigned to Si-O asymmetric stretching vibrations, and another signal in the 900–500 cm−1 band spacing was considered the stretching vibrations of single bond groups.In the FTIR spectrum of the pristine concrete (CC), as demonstrated in brations . The ban- groups ,95,96. T(nitryl) . The peaN- bonds . In the In the FTIR spectrum of C-4D (-bias cross-section), C-4D (+bias cross-section), C-4D (filling cross-section), and C-4D (warp cross-section) as exhibited in Various fractured multiaxis 3D carbon fiber concretes including C-1D-, C-2D-, and C-4D-, as well as the neat concretes on the top, bottom, and cross-sections and their failure modes after the off-axis flexural test are exhibited in The top, bottom, front and back views of the fractured concrete and their failure modes after the off-axis flexural test are illustrated in As shown in Probably, the fracture mechanism comprised regional matrix fracture, matrix-filament bundle debonding for each yarn set in the concrete, filament TOW bridging due to tensile strengthening via principal fibers (0°), partly complex shearing via particularly angular yarn orientations, and shortly after outer partial or complete filament TOWs pull-out and stick-slip based on telescopic movement principles between the outer filament-matrix or core filament-filament frictions. As a result, all yarns carried the flexural load through complex load distribution of the cementitious matrix, such as the warp, filling, and ± bias yarns. Nevertheless, more research is required to define the complex crack control contribution, including some types of stress concentration in the through-the-thickness of the fiber concrete considering each fiber set during flexural loading.Multiaxis 3D carbon/cementitious fiber concretes which were developed for the study, can find several applications such as prefabricated panels, bridge decks, industrial floors, and airport runways. In addition, they can be used for rehabilitation purposes, such as repairing or retrofitting in aging infrastructure applications. Some of the examples of such applications are seismic retrofit, confinement, connectors for numerous settlements, including industrial and nuclear plants and parking structures, bridge decks, and seawall construction . This isMultiaxis three-dimensional polymeric PAN carbon fiber/cementitious matrix concretes were developed. The flexural strength of the C-1D-(0°), C-2D-, and C-4D-(0° and ±45°) concrete composites were considerably higher compared to the pristine concrete. The flexural strength performances for various developed polymeric carbon fiber concrete composites were generally proportional to the fiber orientation and amount of fiber in a particular direction, depending upon the structural fiber architecture. On the other hand, the flexural energy absorption performance of the C-1D-(0°), C-2D-(0° and 90°), and C-4D-(0 and ±45°) concrete composites were outstandingly superior to the neat concrete.Neat concrete exhibited fatal brittle crack failure behavior. On the contrary, the fracture mechanisms in the four directional carbon fiber/cementitious matrix concretes were described as local brittle matrix breakages and broom like-damage features on the filaments, filament bundle-matrix debonding for each yarn set, multifilament bridging in the TOW (perhaps particle bridging in the matrix) due to tensile strengthening and rapid pulling out of the outer filaments and stick-slip between outer filament-matrix or core filament-filament friction. This likely resulted in many micro-cracks in the C-4D concretes, which controlled the number of cracks, crack width, and crack distance. Therefore, the multiaxis 3D fiber polymeric concrete was contemplated damage-tolerant materials compared to the neat concrete. More research is required in future endeavors to model the in-plane and out-of-plane crack initiation and propagation under angular (off-axis) flexural loading for precise designing such structures.


The dimensions of the slab were 2000 × 2000 × 150 mm. Laboratory tests assessed compressive strength, the modulus of elasticity, splitting tensile strength, and bending tensile strength. Based on approximate data from the 3D deformation surfaces, an evaluation of the load-displacement diagrams for the center of the slab and for the center of eccentricity was performed. In conclusion, an overall evaluation and discussion of the results relies on experiments and the mechanical properties of fiber-reinforced concrete.Research on the interaction between slabs and subsoil involves the field of materials engineering, concrete structures, and geotechnics. In the vast majority of cases, research focuses on only one of these areas, whereas for advanced study and computer simulations, detailed knowledge of the whole task is required. Among the new knowledge and information upon which this article focuses is the evaluation of subsoil stress using specialized pressure cells, along with detailed measurements of the deformation of a fiber-reinforced concrete slab. From a design point of view, this research is focused on the issue of the center of the cross section and the influence of eccentricity. Knowledge in this area is not yet comprehensively available for fiber-reinforced concrete slabs, where 2D deformation sections of the slab and 3D deformation surfaces of the slab are used in experiments. The experimental program includes a centrically and eccentrically loaded slab. These are structural elements that were tested on a specialized device. Both slabs had the same concrete recipe, with a dispersed reinforcement content of 25 kg/m One of the most used building materials is concrete ,2. HowevA variant of concrete where these specifics disappear is so-called fiber-reinforced concrete ,13,14,15From the abovementioned works, the significant potential of fiber-reinforced concrete as a construction material for load-bearing structures is evident ,19,20,213. The experiments in [One of the possible uses of fiber-reinforced concrete in construction is in foundation structures and floors ,24,25,26ments in ,24,25,26Therefore, part of this research focuses on the laboratory testing of fiber-reinforced concrete to determine its mechanical properties, while another part of the research focuses on the experimental testing of fiber-reinforced concrete slabs to evaluate their behavior when exposed to the environmental effects of loading. This research builds on previous experiments on slabs in interaction with subsoil using numerical modeling ,28. The In the case of solved tasks, it is necessary to take into account, among other things, the influence of the subsoil itself and the interactions between the foundation and the subsoil and between the structure and the foundation; we call such a task SSI (soil–structure interaction). For this reason, it is also very important to determine the type of subsoil, e.g., by using a geological profile, where the individual subsoil layers and their mechanical properties, such as stiffness (rigid or flexible subsoil) and cohesion, are visible . It is gThe methodology and experimental program of research area were determined on the basis of current design-code approaches in the field of geotechnics  and concThe presented research, as already mentioned, deals with the determination of the mechanical properties of fiber-reinforced concrete on the basis of laboratory testing and experimental testing of fiber-reinforced concrete slabs. Basic tests to determine the compressive strength and modulus of elasticity, as well as detailed tests to determine tensile strength, have been performed ,43,44,45Experimental testing of fiber-reinforced concrete slabs was carried out at the Faculty of Civil Engineering, VSB—Technical University of Ostrava, Czech Republic, where there is specialized loading equipment called Stand . It is p® 3D 65/60 BG fibers [3. The basic mechanical characteristics of the fibers are given in Transport concrete was used for the production of slabs and laboratory samples, where the recipe and the content of individual components in the concrete mixture are given in G fibers  with a dc, cubef) measuring 150 × 150 × 150 mm and on cylinders  with a diameter of 50 mm and a height of 300 mm. The test scheme is shown in The laboratory program included tests of samples to determine the compressive strength, modulus of elasticity, splitting tensile strength, and bending tensile strength ,43,44,45F is the maximum load at failure of the sample, l is the length of the line of contact, and d is the transverse dimension of the body. Very important properties also include tensile strength, which was determined by the splitting tensile test and the bending tensile strength test. Two sample loading variants were selected for the split tensile test. The first variant was for testing the splitting tensile strength perpendicular to the direction of filling. The second variant was for testing the splitting tensile strength parallel to the direction of filling. The schema is shown in Three-point bending test :(2)fc,fF is the maximum failure load of the sample, L is the span, b and h are the transverse dimensions of the specimen, and e is the distance between the support and the force.Four-point bending test:An alternative for determining tensile strength is a bending tensile test. The scheme and dimensions of the samples for both types of tests are shown in From the test record, evaluations of the load–displacement diagrams for the three-point and four-point test were also performed; the diagrams are shown in In the case of a three-point bending test, the load–displacement curves are very similar. After the initialization of the crack, the load will decrease significantly to about 50–60% of the maximum load. The load gradually decreases with the development and opening of the crack. In the test, the crack was close to the load, i.e., at the point of maximum bending moment. In the four-point bending test, the results differ. In one of the selected tests, there was even tensile strengthening. In a four-point test, the location of the crack is not predefined. A crack can and did form between the loading forces. Gradually, the load decreased, and a crack opened up and increased deformations.c, cylf, and the cubic compressive strength, c, cubef, can be expressed as a mutual ratio of their values. The average value of this coefficient is 0.85, where this is the usual value. The split tensile strength was approximately 1/10 of the cylindrical strength. This is also the usual ratio.Based on all laboratory tests, the mechanical properties of fiber-reinforced concrete were determined, with the average values given in A comparison was performed by conversion to uniaxial tensile strength. It is also possible to determine the value of the uniaxial tensile strength from the splitting tensile test as follows: Furthermore, the functional dependence of the splitting tensile strength (perpendicular to the filling direction) and the bulk density was determined, as shown in It is clear from r is the radius of the load plate (150 mm), a1 and a2 are the constants of the 2nd degree polynomial corresponding to the stress dependence of the plate loading . This soil can be globally classified, according to the already invalid standard CSN 731001 , as clay loading , and pmadef,2E/def,1E ratio. In the case of pre-experiment measurements, the ratio of modulus of deformation averaged 2.8. From the above, it can be assumed that the soil has a relatively large deformation capacity. The modulus of deformation ratio determined after the tests decreased to 2.1. The measurement results also show that def,1E is very similar before and after the tests and is in the range of 16–19 MPa. The differences for the modulus of deformation def,2E are more pronounced, in the range of 32–58 MPa.When setting the modulus of deformation of the subsoil, it is appropriate to determine the To determine the physical and descriptive characteristics of the soil, two samples were taken at different places under the base slab. The strength characteristics were determined on a reconstituted 100% sample on a box shear device (sample size: 100 × 100 × 20 mm).The experimental tests consisted of two fiber-reinforced concrete slabs that differed in the position of the load. The first slab, G10, was loaded in its center, so it was a centrically loaded slab a.In the case of the eccentrically loaded slab G11, the position of the load was situated 400 mm from the center of the slab, i.e., it was an eccentrically loaded slab G11 b. Both fBased on the experimentally measured data, a diagram  was creaThe first part of the experimental measurement of the slabs was monitoring the deformation behavior of the slab in the transverse direction. In the framework of this article, the sections correspond to track sensors 06–22 a, 23|23*The maximum negative deformations in the slabs of track sensors 06–22, which correspond to the lifted corners of the slabs, measured 13.93 mm  and 20.01 mm , respectively. The maximum positive deformations in the slabs of track sensors 06–22, which correspond to the pushing of the slabs into the subsoil, were measured at 26.72 mm  and 8.24 mm , respectively. Comparative 2D deformation sections at the locations of track sensors 06–22, which capture the behavior of both fiber-reinforced concrete slabs, are shown in The maximum negative deformations at the locations of track sensors 23|23*–29|29*, which correspond to the lifted corners of the slabs, were 8.93 mm  and 10.57 mm , respectively. Maximum positive deformations at the locations of track sensors 23|23*–29|29*, which correspond to the pushing of the slabs into the subsoil, were 24.36 mm  and 18.79 mm , respectively. Comparative 2D deformation sections at the locations of track sensors 23|23*–29|29*, which depict the behavior of both fiber-reinforced concrete slabs, are shown in In the longitudinal direction, the section corresponding to track sensors 03–42 was evaluated . From thThe maximum negative deformations in the slabs of track sensors 03–42, which correspond to the lifted corners of the slabs, were 13.06 mm  and 20.37 mm , respectively. The maximum positive deformations in the slabs of track sensors 03–42, which correspond to the pushing of the slabs into the subsoil, were 26.72 mm  and 21.84 mm , respectively. Comparative 2D deformation sections in the slabs of track sensors 03–42, which capture the behavior of both fiber-reinforced concrete slabs, are shown in Centrically loaded slab G10 and eccentrically loaded slab G11 during loading are shown in The 3D deformation surfaces of centrically loaded slab G10 and eccentrically loaded slab G11 were also evaluated for selected loading steps, i.e., for loading steps of 225 and 450 kN. Part of the evaluation of 3D deformation surfaces was the evaluation of the 3D deformation surface with transfer into the projection, where the maximum deformation area is clearly visible with regard to the applied load and eccentricity. The projections are shown only for the central part of the slab, i.e., for a slab strip of size 1 m. The evaluations of the 3D deformation surfaces of centrically loaded slab G10 and eccentrically loaded slab G11 and the projections of the 3D deformation surfaces for a loading step of 225 kN are shown in The evaluations of the 3D deformation surfaces of centrically loaded slab G10 and eccentrically loaded slab G11 and the projections of the 3D deformation surfaces for a loading step of 450 kN are shown in Based on the approximation of the 3D deformation surfaces of the slabs, load–displacement diagrams were evaluated, as shown in Based on the measured pressures in the subsoil during the load test of the slabs, the stresses in three layers at the locations of the stored pressure cells were evaluated. A diagram of the arrangement of the pressure cells is shown in The individual layers of the pressure cells or the stress curves in the pressure cells are differentiated by color in It can be seen from 3 is suitable with regard to concreting technology and financial costs. However, in comparison with previous experiments, it can be stated that the total load-bearing capacity of the fiber-reinforced concrete slab  will typically be less than the traditional solution of a reinforced concrete slab with concrete reinforcement, where the load is up to 700 kN. However, the use of concrete reinforcement is more demanding on concreting and overall financial costs. However, the use of only plain concrete for slabs has the effect of creating a significantly lower load-bearing capacity of the slab than using a fiber-reinforced concrete slab. The collapse of a plain concrete slab is also marked by a significantly quasi-brittle nature.The laboratory tests and tests of structural elements verified that fiber-reinforced concrete can significantly contribute to the structural and material optimization of the design of concrete structures. The use of fiber-reinforced concrete can significantly simplify the production of concrete structures. The selected dosing of fiber of 25 kg/mThe use of fiber-reinforced concrete also makes a significant contribution to increasing the ductility of the structure, i.e., to the safety of the structure as compared to the use of plain concrete. A broader evaluation of the presented experiments and tests already performed revealed that the overall load-bearing capacity of the slab in the subsoil interaction is influenced by the mechanical properties of the concrete rather than the subsoil. The mechanical parameters of the subsoil significantly affect the resulting deformations of the slab. These can be differences of up to tens of percentage points. From a design point of view, the experiment performed on a fiber-reinforced concrete slab with an eccentric load showed that the resulting load-bearing capacity was very similar to that of a slab with a centric load. This was mainly due to the fact that the slab remained whole throughout the test. However, the eccentric load significantly affected the course of deformations.The selected fiber-reinforced concrete variant has a number of advantages over plain concrete. However, the quasi-brittle nature of fiber-reinforced concrete in itself requires a set of laboratory tests of the compressive strength of concrete, tensile strength under splitting, modulus of elasticity, and bending tensile strength. From the results of laboratory tests and a comparison with the available recommendations, we confirmed that it is appropriate to carry out specialized tests rather than using only the recommendations for modulus of elasticity and tensile strength under splitting.It is clear from the tensile tests that the strengths different may be for test setup. However, the recalculated tensile strengths are relatively similar. In the three-point test, the strength values are slightly higher than in the four-point bending test. This is because the location of the crack is not predetermined in the four-point bending test. In the case of the modulus of elasticity, the results of the tests show a significantly lower modulus of elasticity than in Model Code 1990 [c (kPa), and angle of internal friction, φ (°) 34.4. It is suitable to perform tests on the deformation modulus of soil, where the measurement values a relatively wide variance of measured values from 16 to 58 MPa; the resulting values depend on the actual history of loading and the specific climatic conditions. The mechanical properties of the soil are also often affected by the climate at the surface in particular. It is advisable to repeat this test several times to obtain a wider set of static data.Important information also includes knowledge about the subsoil, where our analysis revealed that the subsoil has the character of incoherent soil . According to the percentage of individual fractions, this soil can be globally classified as clayey sand. The basic soil parameters are cohesion, Simplified analytical/numerical models and experiments, such as symmetry, dimension, etc., can be used for typical cases of a slab in interaction with subsoil. Important aspects of the problem include the need to closely monitor the behavior and parameters of the fiber-reinforced slab and subsoil. In the case of fiber-reinforced concrete slabs, a network of vertical deformation sensors was used. In the case of subsoil, specialized pressure cells were used to measure stress. The comparison of the results allows for a detailed view of the behavior of the fiber-reinforced concrete slabs during a load test. A comparison of the results from the fiber-reinforced concrete slab and the subsoil showed that the initial formation of cracks at the lower surface of the slab were manifested by an increase in the contact stress in the middle of the slab. There was a significant increase in stress in the pressure cell in the middle of the slab closest to the surface. A more detailed study of data from soil pressure sensors in the lower layer (at a depth of 0.8 m) revealed that the stresses are significantly lower compared to the contact stresses at the surface. By approximating the stress results in the pressure cells in the individual layers, it was found that the concerned area of the subsoil and the deformation zone for maximum load is about 2 m deep. The specific values measured at the edges of the slab were significantly lower than below the center of the slab. The differences between the measured values were significant. With the achievement of the maximum load-bearing capacity of the fiber-reinforced concrete slab and the formation of significant cracks in the concrete, the stress in the pressure cells no longer increased and remained at the maximum values. This was due to the fact that there was a significant decrease in the bending stiffness of the slab. The differences in pressure cells in the subsoil for the centrically (G10) and eccentrically (G11) loaded slabs are relatively small. Only in the case of higher loading steps, when the corners of the slab were already being lifted, was it evident that the pressure cell was relieved, while the character itself was slightly different for the centrically (G10) and eccentrically (G11) loaded slabs. This is due to the different geometry and position of the load.Research on slabs in interaction with subsoil is very demanding. This is because a number of factors enter into the analysis, which involves the use of advanced testing and diagnostic methods. The selected variant of fiber-reinforced concrete slab has a number of advantages in comparison with concrete and reinforced concrete slabs. However, the quasi-brittle nature of fiber-reinforced concrete itself requires an extensive set of laboratory tests. From the point of view of the design of concrete structures, it is important to monitor the deformations and cracks in the concrete of the fiber-reinforced concrete slab itself. By a more detailed study of the results from 2D deformation sections, it is possible to distinguish when the increase in deformations is no longer elastic and cracks have formed in the concrete. It is also clear from the graph that the deformation increment gradually increases within the loading step and the bending stiffness of the fiber-reinforced concrete slab decreases. The fiber-reinforced concrete slab remained whole in both tests; when the maximum load was reached, the cracks were over the entire thickness of the slab. During the test, the edges of the slab gradually began to rise. When analyzing the deformations of the slab with centric and eccentric load, it is also possible to use 3D graphs, which enable a clearer and more comparable display of deformations.The differences in the overall load capacity of slabs with centric and eccentric loads are relatively small. The center of the eccentric load was already outside the core of the cross section.For lower load steps, the difference in fiber-reinforced-concrete-slab behavior and deformation was relatively small. The differences began to increase, especially in the second half of the load experiment tests (loading).The position of the steel load plate affected the position of the cracks in the slab. For the slab with an eccentric load, the deformations were more pronounced on the load side. In the case of a slab with an eccentric load, there was also a more pronounced lifting of the corner on one side of the slab.In comparison with previous experiments, it can be stated that better deformation parameters of the subsoil mainly affected the magnitude of deformations of the slab and the subsoil rather than the overall load-bearing capacity.It is advisable to repeat the test several times when determining the deformation modulus of the subsoil and the ratio for the secondary and primary cycle, which are used to determine the deformation capacity of the subsoil. Pressure cells in at least two levels are suitable for determining the active deformation zone of the subsoil. By far, the most important are the pressure cells directly under the load in plan view.3 can significantly contribute to increasing the overall load-bearing capacity and eliminating cracks in the concrete.The laboratory experiments and tests of structural elements confirmed that dosing fibers at 25 kg/mThe use of fiber-reinforced concrete requires a more detailed determination of its mechanical properties. Tensile strength is especially important. It is also important to produce load–displacement diagrams, which can be used to identify fracture mechanical parameters.From laboratory tests in split tension, it was found that the position of concreting affects the mechanical properties in tension, i.e., the number of fibers along the cross-sectional height varies slightly.It is important to perform tests of the modulus of elasticity of concrete, where the values found are significantly different from those in the recommendations. Modulus of elasticity values were lower than those of the recommendations.When evaluating slab tests in subsoil interaction, it is appropriate to use 3D deformation graphs and to determine the load–displacement diagram for the entire structural element tests.The results of the research can be presented in partial conclusions:The benefits of this research include that a comprehensive set of information and knowledge with a detailed evaluation of laboratory tests and tests of structural elements can be further used for advanced numerical-modeling-based non-linear analysis, where an optimized structure solution is solved.Among the important and new knowledge of the presented research is that the problem of the slab in the interaction of the subsoil can be further extended by the structural solution of the sliding joint. This area can be followed by laboratory research into the properties of sliding joints in the climate chamber and measurements on real structures of the base slab with a sliding joint on the subsoil. The authors will focus on this area in further research.


This study was performed to appraise the biocompatibility of polyhedral oligomeric silsesquioxane (POSS)-grafted polyurethane (PU) nanocomposites as potential materials for muscle tissue renewal. POSS nanoparticles demonstrate effectual nucleation and cause noteworthy enhancement in mechanical and thermal steadiness as well as biocompatibility of resultant composites. Electrospun, well-aligned, POSS-grafted PU nanofibers were prepared. Physicochemical investigation was conducted using several experimental techniques, including scanning electron microscopy, energy dispersive X-ray spectroscopy, electron probe microanalysis, Fourier transform infrared spectroscopy, and X-ray diffraction pattern. Adding POSS molecules to PU did not influence the processability and morphology of the nanocomposite; however, we observed an obvious mean reduction in fiber diameter, which amplified specific areas of the POSS-grafted PU. Prospective biomedical uses of nanocomposite were also appraised for myoblast cell differentiation in vitro. Little is known about C2C12 cellular responses to PU, and there is no information regarding their interaction with POSS-grafted PU. The antimicrobial potential, anchorage, proliferation, communication, and differentiation of C2C12 on PU and POSS-grafted PU were investigated in this study. In conclusion, preliminary nanocomposites depicted superior cell adhesion due to the elevated free energy of POSS molecules and anti-inflammatory potential. These nanofibers were non-hazardous, and, as such, biomimetic scaffolds show high potential for cellular studies and muscle regeneration. Hybrid (organic/inorganic) nanocomposites prepared from polyhedral oligomeric silsesquioxane (POSS) have drawn substantial attention due to numerous beneficial purposes that typically originate from distinctive physical properties of POSS. Tissue engineering is a very promising field that offers surprising ventures in regenerative medicine. This area is focused on the search for substitutes to alleviate pain by reinstating the functions of injured cells and organs . TherefoIn this report, we portray the successful synthesis of a polyhedral oligomeric silsesquioxane (POSS)-grafted polyurethane (PU) blend of nanofibers via a facile electrospinning route. We chose the PU polymer and mouse myoblast cell lines C2C12) for our study, because PU has marketable elastomers, which can allow for its use in adhesives, flexible films, coatings, and hard plastics ,12, and  for our Keeping in consideration the distinctive characteristics of PU and POSS, we dispersed 10 wt.% of POSS molecules in a PU solution. A multipurpose and economical electrospinning method was utilized to spin the PU/POSS composite solution, and it is typically used to generate elongated and incessant fibers ,23 for aN,N-dimethylformamide (DMF)  were utilized as solvents devoid of distillation. C2C12 cells and CCK-8 were obtained from American Type Culture Collection (ATCC) and Sigma–Aldrich Chemical Co., Burlington, MA, USA. Dulbecco’s modified Eagle medium (DMEM) was obtained from Gibco® life technologies, Grand Island, NY, USA.Polyurethane  was obtained from Cardio Tech. Intern., Tokyo, Japan. Isocyanatopropyl dimethyl silylcyclohexyl-polyhedral oligosilsesquioxane (POSS macromer) was a kind gift from Air Force Research Lab, Montgomery, OH, USA. Tetrahydrofuran (THF) and w/w) with the help of magnetic shaking without heating. The POSS (10 wt.%) in THF was supplemented with PU (wt. ratio 10:1). The PU–POSS hybrid mixture was vigorously stimulated for 3 h without heating for appropriate assimilation. The PU–POSS hybrid mixture was transported into a 10 mL syringe to synthesize a POSS-grafted PU nanofibrous matrix. The electrospinning was conducted at a voltage of 12 kV and at a remoteness of 15 cm between the needle and collector. The pristine and blend nanofibrous scaffolds were dried in an oven at 60 °C for one day.Fabrication of pure and POSS-loaded PU nanofibrous scaffolds were conceded using a simple electrospinning approach. PU (10 wt.%) was obtained through liquefying PU pellets in DMF: THF  was applied to trace out the exterior morphology of the PU and POSS-grafted PU mats at different magnifications. The dried pristine (PU mat) and blend (POSS/PU) scaffolds were dispersed on carbon tape, layered with platinum by a sputter coater for 10 s at a current of 70 µA and a voltage of 15 kV. POSS-grafted PU nanofibers were screened for chemical concerto using an energy dispersive X-ray (EDX) spectrometer . The allocation of elements was meditated by electron probe microanalysis (EPMA). Fiber diameter was unswervingly considered from SEM metaphors. In all cases, 25 to 30 fibers from a randomly chosen section were calculated, and the mean was estimated. Fiber dimension distribution was measured using ImageJ (Scriptable Java app for scientific image processing) software. The XRD of PU and POSS-grafted PU blend nanofibrous scaffolds were documented on an X-ray diffractometer  with copper Kα radiation (λ = 1.540 Å) over Bragg angles ranging from 10 to 60. The voltage and current of the apparatus were fixed at 30 kV and 40 mA, respectively. The effect on the polyurethane polymer structure after integration of POSS nanocrystals was examined by Fourier transforms infrared spectroscopy . The spectra of PU and POSS-grafted PU were verified by pulverizing respective mats with KBr to obtain pellets that were examined by adjusting the wave number from 4000 to 400 cm2 culture flasks . The C2C12 cells were revived in DMEM (pH 7.4) including 10% fetal bovine serum (FBS) and 1% penicillin–streptomycin. The cultivated C2C12 cells were sited in a humidified incubator and were provided with 5% CO2, 95% air, and a temperature of 37 °C.The muscle cell–scaffold interactions on the designed POSS-grafted PU blend nanofibers were considered by culturing C2C12 in 75 cmThe POSS-grafted PU nanofibrous scaffolds were cut into 1 × 1 cm dimensions, sterilized, and kept in the DMEM medium as described previously . The POSThe viability and vitality of C2C12 cells cultured on POSS-grafted PU scaffolds were scanned by using quantitative CCK-8 assay, which was performed spectrophotometrically. In brief, C2C12 cells were replenished with fresh media at 1, 3, and 5 days, and 10 μL of water-soluble tetrazolium-8 (CCK-8) was supplemented in individual wells and incubated. The results were obtained by considering absorbance at 450 nm by a spectrophotometer  as previously described .4 cells were cultured in culture dishes containing DMEM to gain 70–80% convergence. In due course, the exhausted media was reinstated with differentiation medium (DMEM with 5% (v/v) horse serum, 2 mM glutamine, penicillin (100 U/mL), and streptomycin (100 lg/mL)). The cultured myoblasts were maintained in differentiation medium until the termination of analysis (0–7 days). The formation of myotube was screened on a regular basis with a microscope.C2C12 differentiation was studied on POSS-grafted PU nanoscaffolds as articulated earlier . In briev/v) in 1 mM of sodium phosphate buffer (pH 5.6). Accordingly, 1 mL was dispensed among cultivated cells and incubated for about 5 min. The myoblasts were washed off using distilled water. Later, immobilized cells were marked with Giemsa stain (1 mL) prepared as 1:20 in 1 mM of sodium phosphate buffer and developed at room temperature for 10 min. Finally, C2C12 myoblasts were once more cleaned using DW. To end the process, the stained cells were captured from different angles with a camera equipped with an inverted microscope. The progress and expansion of myotubes were visualized using intense coloration with prominent nuclei. The unstained immobilized culture was sealed and stored at RT until future use.The cultured myoblasts on POSS-grafted PU nanoscaffolds were bathed in phosphate-buffered saline, immobilized in absolute methanol approximately for 2–5 min, and subsequently air dried for 10 min. The Jenner staining  was diluted as 1:3 (Staphylococcus aureus (ATCC strain no.29213) was cultivated to evaluate antibacterial activity of POSS-grafted PU nanoscaffolds as previously reported [S. aureus culture was geared up on nutrient agar . Individual colonies were picked up and preserved in cryovials at freezing temperature (−80 °C) in anticipation of further utilization. The S. aureus was grown (106 CFU) with POSS-grafted PU nanoscaffolds (varying concentrations) to investigate the lowest amount needed to restrain biofilm formation. The established growth parameters were temperature: 37 °C, rpm: 150.In this study, reported . The S. The developments in tissue engineering have drawn incredible awareness to the field. The reason for this constantly growing demand is a result of its projected medical relevance for issues such as the renewal and rejuvenation of smashed tissues and organs. Currently, tissue engineering appears to be the most competent and noteworthy move that endeavors to present solutions for organ transplantation as well heralding innovative perspectives for management of diverse syndromes.In this study, we optimized POSS concentration for successful preparation of nano-organic–inorganic hybrid (PU–POSS) scaffolds via a facile electrospinning route and investigated their response to C2C12 cells. The presence of dispersed POSS nanocrystals in hybrid matrices was also confirmed by XRD . Due to Additionally, incorporation of POSS into PU was also established by EPMA . In the −1, which are assigned to ν(N–H), ν(C–H), ν(C=O), ν(C=C), ν(C–C), ν(C–O), and ν(C–H), respectively, on substituted benzene [To further investigate the interaction between POSS nanocrystals an PU, FT-IR analysis was performed. The infrared spectrum of PU and POSS-grafted PU is illustrated in  benzene ,37 (FiguAfter optimizing the weight percentage of POSS in the PU solution and the successful fabrication of POSS-grafted PU scaffold, cell–nanofiber interactions of pure PU and POSS-grafted PU scaffolds in this study were further investigated. The POSS-grafted PU scaffolds were scrutinized for biocompatibility with C2C12 cells. The C2C12 cells were seeded on PU and POSS-grafted PU scaffolds, and, correspondingly, biocompatibility assay was carried out for one, two, and three days. Concerning the POSS-grafted PU nanomatrix, the quantum of healthy cells exponentially amplified. In the case of PU, no toxicity was found, but a higher number of live cells were on POSS-grafted PU nanofibers a. It wasS. aureus depicted exponential doubling during the timeframe between 4 to 8 h in a virgin environment (Additionally, a simplistic way to estimate myotube development was designed in this study. ironment d. The enironment . AdmitteIn conclusion, nanoscaffolds with superior biocompatibility as well as proper potency to retain their form can be used for biointegration and will unquestionably provide a priceless substitute to current materials. Our study highlights the benefits of POSS grafting for C2C12 cellular biocompatibility. The POSS-grafted PU scaffolds were fabricated by electrospinning. The SEM-EDX, EPMA, and XRD analyses clearly confirmed the formation of a POSS-grafted PU nanoscaffold. In the present study, it was established that the addition of POSS in PU-shaped grafted fibers had a smaller width and lesser diameter distribution than those of electrospun PU nanofibrous scaffolds. Furthermore, it was noted that cell interaction and dispersal behavior were more expressive in the case of POSS-grafted PU nanofibers. We hypothesize that POSS-grafted PU scaffolds possess a superior surface-free energy of POSS siloxanes; therefore, they are responsible for fine attachment and proliferation. The POSS-grafted PU nanoscaffold will act as a novel contrivance for nanomedicine, tissue engineering, and C2C12 differentiation.