The safe disposal of an enormous amount of waste glass (WG) in several countries has become a severe environmental issue. In contrast, concrete production consumes a large amount of natural resources and contributes to environmental greenhouse gas emissions. It is widely known that many kinds of waste may be utilized rather than raw materials in the field of construction materials. However, for the wide use of waste in building construction, it is necessary to ensure that the characteristics of the resulting building materials are appropriate. Recycled glass waste is one of the most attractive waste materials that can be used to create sustainable concrete compounds. Therefore, researchers focus on the production of concrete and cement mortar by utilizing waste glass as an aggregate or as a pozzolanic material. In this article, the literature discussing the use of recycled glass waste in concrete as a partial or complete replacement for aggregates has been reviewed by focusing on the effect of recycled glass waste on the fresh and mechanical properties of concrete. Glass is one of the world’s most diverse substances because of its substantial properties, such as chemical inertness, optical clarity, low permeability, and high authentic strength ,2,3. TheRecycling and reducing waste are key parts of a waste-management system since they contribute to conserving natural resources, reducing requests for waste landfill space, and reducing pollution of water and air ,18. AccoIn fact, innovative options for recycling WG must be developed. One significant option is to use WG for construction materials . The recBesides the above-mentioned dangers of WG and the need to recycle it economically and environmentally, this research explores the source of WG as well as its physical and chemical characteristics. In addition, this study aims to review the literature that discusses the use of recycled WG in concrete as a partial or complete alternative to aggregates by focusing on the effect of this waste on the fresh and mechanical properties of concrete in order to demonstrate the possibilities of using recycled WG in concrete and to provide practical and brief guidance. Furthermore, we are establishing a foundation for future study on this material and describing research insights, existing gaps, and future research goals.Glass exists in various colors and types, with various chemical components. The physical and mechanical properties of crushed WG are listed in There are two parallel points of view on the workability of WG-containing concrete. A review of past studies on the impact of WG aggregates on the mixes of workability is summarized in Contrastingly, some studies have stated that including waste glass into the mixes lowered workability. Nevertheless, such a decrease has been associated with sharp edges, higher glass particle aspect ratio, and angular form, with obstruction of the movement of particles and cement mortar ,69,70,71Past studies on the impact of WG aggregates on the bulk density, which are summarized in 3) was larger than that of fine aggregate (2574 kg/m3) [On the other hand, Liu, Wei, Zou, Zhou and Jian  stated t4 kg/m3) ,103,104.By reviewing past studies on the impact of WG aggregates on the compressive strength of waste-glass concrete, summarized in In order to better understand the impact of glass waste on the properties of the waste-glass concrete ,113,114.However, some studies have stated that waste glass increases mechanical strength. This increase is primarily realized because of the surface texture and strength of the waste glass particles compared to natural sand ,117,118 Regarding the influence of WG color on properties, some studies have stated that the color of WG did not produce any noticeable variation in strength ,122. On Past studies on the impact of WG aggregates on the splitting tensile strength of waste-glass concrete, which are summarized in In contrast, Jiao, Zhang, Guo, Zhang, Ning and Liu  indicateThe flexural strength of waste-glass concrete shows comparable tendencies to its compressive strength and tensile strength. Most of the research revealed that introducing WG aggregates reduced flexural strength. However, other research showed that flexural strength increased when WG was included ,135,136.Moreover, it can be concluded that the discrepancy between studies may be related to the type, size, and source of WG used in the mixtures. The mineral composition varies as the type of glass changes. Therefore, changing the mechanisms of interaction with binders in concrete, in turn, affects the properties. The modulus of elasticity of concrete (MOE) depends on the normal and lightweight aggregates elasticity modulus, cement matrix, and their relative ratios in the mixes . In geneHowever, some studies have stated that including WG decreases the MOE of concrete. For instance, Ali and Al-Tersawy  substituThe workability of waste-glass-containing concrete mixtures for fine or coarse aggregates was less than for natural aggregate-containing mixtures. Nevertheless, despite the poorer workability, some studies found that the mixtures were still workable.Most studies indicated that with the introduction of WG, the density of concrete decreased due to the decreased density and specific gravity of waste glass aggregates.The findings of the literature have been somewhat indecisive regarding the properties of concrete, such as compressive strength, splitting tensile strength, flexural strength, and modulus of elasticity.The findings revealed that the compressive strength, splitting tensile strength, and flexural strength of concrete deteriorated by integrating WG. Nevertheless, the findings concerning the elastic modulus of concrete were conflicting. This decrease was essential because of the sharp edges and smooth surface of the waste glass that caused the poorer bond between cement mortar and waste glass particles at the ITZ.Studies also showed that the optimal aggregate substitution level was about 20%. In addition, the glass color does not have a substantial influence on the strength. Although the results are indecisive, WG has the possibility to be an acceptable substitute for fine or coarse concrete aggregates in concrete.Adding waste glass to the concrete mixture may improve certain mechanical characteristics of concrete, reduce concrete dead load, and provide an ecological substitute for normal aggregates.The utilization of WG in concrete affects the fresh and mechanical properties of waste-glass concrete, which must be taken into consideration before being used in structures. The overall conclusions of this review are:More investigation is required into the mechanical characteristics of high-performance and high-strength waste-glass concrete.The effects of different glass kinds and colors on concrete mixes should be thoroughly investigated in the future.Test fewer common types of glass as aggregates in concrete because the vast majority of research only covers soda-lime glass.Conduct a comprehensive evaluation of the real environmental effects through life-cycle assessment to evaluate the feasibility of using this waste.This paper makes the following broad recommendations for future investigations:


Background: Nationally, a decrease in total antibiotic use in nursing homes during the COVID-19 pandemic was observed with an increase in select agents used for respiratory infections. Currently there is minimal data on antibiotic use in long-term care facilities (LTCFs) in Tennessee. To address this issue, the Tennessee Department of Health (TDH) developed a monthly point-prevalence survey of antibiotic use. Utilizing this tool, we sought to determine the effect the pandemic had on antibiotic use in Tennessee LTCFs. Method: We developed a REDCap questionnaire to collect information on selected antibiotics administered in Tennessee LTCFs. Antibiotic use percentage was determined by dividing the number of residents who received an antibiotic on the day of survey by facilities’ average censuses. Data were divided into a prepandemic period (January 2019–February 2020) and a period during the pandemic (March 2020–December 2021). Antibiotic prescriptions were grouped into 4 classes according to their most common uses: Clostridium difficile infections, urinary tract infections, skin and soft-tissue infections (SSTIs), and respiratory infections. Average percentage of residents on antibiotics were compared between study periods. Results: In total, 37 facilities participated in the survey during the prepandemic period and 32 facilities participated during the pandemic period; 14 participated during both periods. The average percentage of residents on antimicrobials before the pandemic was 16.3%, which decreased to 11.5% during the pandemic period (P = .04). During the prepandemic period, 40.2% of antibiotics prescribed were in the common for SSTI category and 38.3% were in the common for respiratory infections category (P = .01); during the pandemic period, 64.3% of antibiotics prescribed were in the common for SSTI category and 45.8% were in the common for respiratory infections category (P = .01). The 3 most prescribed antibiotics in the prepandemic period were amoxicillin (148 prescriptions), doxycycline (140 prescriptions), and levofloxacin (135 prescriptions). The 3 most prescribed antibiotics during the pandemic were doxycycline (141 prescriptions), levofloxacin (125 prescriptions), and trimethoprim–sulfamethoxazole (115 prescriptions). Conclusions: Survey results revealed that antibiotic prescriptions commonly used for respiratory infections increased 7.5% during the pandemic study period. Additionally, the average percentage of residents on antimicrobials fell 4.8% during this period. Both statistics reflect what has been seen nationally with a decrease in antibiotic use with an increase in respiratory antibiotics. This could be due to multiple factors including decreased reporting, a change in healthcare delivery during the pandemic, and facilities seeing an increase of respiratory tract infections. These data will be used to guide future TDH antibiotic stewardship efforts in the long-term care setting.Funding: NoneDisclosures: None


Utilization of additive manufacturing (AM) is widespread in many industries due to its unique capabilities. These material extrusion methods have been developed extensively for manufacturing polymer and polymer composite materials. The raw material in filament form are liquefied in the liquefier section and are consequently extruded and deposited onto the bed platform. The designed parts are manufactured layer by layer. Therefore, there is a gradient of temperature due to the existence of the cyclic reheating related to each deposited layer by the newer deposited ones. Thus, the stated temperature evolution will have a significant role on the rheological behavior of the materials during this manufacturing process. Furthermore, each processing parameter can affect this cyclic temperature profile. In this study, different processing parameters concerning the manufacturing process of polymer and polymer composite samples have been evaluated according to their cyclic temperature profiles. In addition, the manufactured parts by the additive manufacturing process (the extrusion method) can behave differences compared to the manufactured parts by conventional methods. Accordingly, we attempted to experimentally investigate the rheological behavior of the manufactured parts after the manufacturing process. Thus the three-point bending fatigue and the tensile behavior of the manufactured samples were studied. Accordingly, the effect of the reinforcement existence and its direction and density on the tensile behavior of the manufactured samples were studied. Therefore, this study is helpful for manufacturers and designers to understand the behaviors of the materials during the FFF process and subsequently the behaviors of the manufactured parts as a function of the different processing parameters. Additive manufacturing (AM) is growing due to the low amount of wasted material in the manufacturing process and the ability to manufacture complex shapes ,2,3. TheChristiyan et al.  investigFew studies have been performed to investigate the effect of the utilized fiber reinforcements on the rheological behavior of the materials during and after the manufacturing process. This study tried to investigate the rheological behavior of the materials during and after the FFF process by considering the role of the fiber reinforcements. The selected materials for this study were PA6, short carbon fiber-reinforced PA6 (CF-PA6), and continuous glass fiber-reinforced CF-PA6 composite materials.®. The chopped carbon fibers had a mass content of 6.5% in the CF-PA6 filament based on the pyrolysis process. The characteristics of the utilized filaments as the raw materials are presented in The selected materials for this study were polyamide 6 (PA6) and short carbon fiber-reinforced polyamide 6 (Onyx or CF-PA6) produced by MarkForgedAs for the investigation of the fiber-reinforcement impact and the processing parameter effects on the rheological behavior of the materials during the FFF process, a single wall layer specimen  was desiAn Optris PI450 infrared camera was applied in the conducted study of the processing parameter impacts on the thermal and mechanical properties of the polymer and polymer-based composites using the FFF process. The stated infrared camera was positioned at a specific predetermined distance from the extrusion location of the printer. This attempt aimed to monitor and observe a consistent plain field of view (FOV) across all consecutive layers. The temperature rise during the performed fatigue test was monitored and measured by the stated infrared camera as well. Regarding some technical specifications of the infrared camera used in this experiment, frequency, accuracy values, wavelength range, optical resolution, and the rate of the frames were 32 Hz, 2%, 8–14 μm, 382 × 288 pixels, and 80 Hz, respectively. Using a scanning electronic microscope , observations and image analyses were performed to qualitatively analyze the composite microstructure, especially in relation to damage assessment. In this study, an optical microscope  was used to assess the quality of the manufactured samples at the various selected processing parameters with magnifications from 100 to 500 mm.g), crystallization temperatures (Tc), and their heat capacities based on the selected processing parameters. DSC characterization was performed on the raw filament materials (CF-PA6 and PA6) over three temperature ramps: range of 20 to 220 °C under a rate of 10 °C/min. This attempt caused an elimination of the thermal history of the filaments concerning their production process. Furthermore, the printed and fabricated specimens were analyzed by DSC in two ramps (heating and cooling).By means of differential scanning calorimetry (Q1000), both raw filament materials and printed specimens could be assessed for their respective glass transition . The considered strain ratio was 0.1 (Rσ = 0.1); moreover, the related mean strain level was 0.55 εmax. This test was conducted on the short carbon fiber-reinforced polyamide 6 (Onyx or CF-PA6) and the continuous glass fiber-reinforced polyamide 6 (Onyx + GF) composite specimens to be able to study the effect of the continuous glass fiber reinforcement and its subsequent impact on the rheological behavior of the FFF-processed specimens. During the three-point bending fatigue test, the temperature of the specimens was raised, which was measured by the infrared camera to determine the importance and influence of the utilized continuous glass fiber reinforcement in this phenomenon. The dimensions of the prepared rectangular composite specimens were 120 × 10 × 4 mm3. In order to perform a fatigue test, at least three specimens were prepared for each condition.As for the fatigue test of the fabricated specimens concerning the study of the rheological behavior of the FFF-processed composite materials (after the manufacturing process), a three-points bending fatigue test was conducted at various applied maximum strains , liquefier temperature , layer height , and bed platform temperature . The aforementioned processing parameters were studied individually by considering the rest of the processing parameters as constant. The reference values concerning each of the selected parameters were 15 mm/s as print speed, 240 °C as liquefier temperature, 0.1 mm as layer height, and 25 °C as bed temperature. The selected materials for this section were PA6 and CF-PA6 [In light of the importance of the print speed parameter in production time, it was studied to how it impacts the rheological behaviors. To study the impact of this manufacturing processing parameter, 13, 15, and 17 mm/s as the different print speed values were taken into consideration. According to The impact of the liquefier temperature on the rheological behavior of the PA6 and CF-PA6 specimens were investigated. The decided liquefier temperature values for this study were 240, 230, and 220 °C. Regarding the effect of the short fiber reinforcement and the consequent differences between the FFF-processed polymer and polymer composite material, firstly, the crystallinity percentage values were highlighted. Based on the obtained results from the DSC characterization of the FFF-processed CF-PA6 and PA6 samples, the crystallinity percentage values of the CF-PA6 specimens were higher than the PA6 specimens manufactured under the same processing parameters. The chopped carbon fibers in CF-PA6 were the superior sites and spots for the nucleation of the crystalline sections. Furthermore, another difference between the FFF-processed polymers and polymer composite materials is their tensile strength. Regarding compression between the obtained tensile strength values of CF-PA6 and PA6, it was understood that the single-wall FFF-manufactured CF-PA6 specimens exhibited lower strength. The short carbon fibers as the solid component in the molten polymer decreased the fluidity of the CF-PA6 during the FFF process compared with PA6 material. The stated decrease in the fluidity during the manufacturing process can be illustrated by the narrow width zones in the manufactured specimens . The staCF-PA6 and PA6 specimens were investigated for their rheological behavior influenced by the layer height. The investigation of the effect of the short fiber reinforcement on the rheological behavior of the materials was carried out by selecting the three different layer height values of 0.1, 0.2, and 0.3 mm, during and after the FFF process. The tensile strength values of 55 ± 0.6, 49 ± 2.5, and 56 ± 2.5 MPa concerning the manufactured CF-PA6 specimens by the layer height values of 0.1, 0.2, and 0.3 mm were achieved, respectively. For better investigation of the effects of the layer height and the short carbon fiber reinforcements on the rheological behavior of the material during the FFF process, a thermal camera for the in situ temperature monitoring was utilized. By considering the obtained results from the tensile test and the in situ temperature monitoring during the FFF process and their subsequent correlations, the existence of the two effective phenomena was declared. The stated two phenomena were firstly the fluidity decrease in the printed materials and secondly the residual/retained temperature increase in the deposited layers during the cyclic raised temperature during the FFF process. The peaks (upper limits) of the obtained time–temperature curves during the manufacturing process revealed the first-stated phenomenon, and the lower limits of these curves revealed the second-stated phenomenon. By increasing the layer height during the manufacturing process from 0.1 to 0.2 mm, the first phenomenon was more highlighted and effective. The decreased fluidity was found according to the decrease in the measured temperature profile peaks by increasing the selected layer height from 0.1 to 0.2 mm a. This dAs for the study of the influence of the last selected processing parameter, the importance of the bed platform temperature on the manufactured specimens was evaluated. The studied materials were PA6 and CF-PA6. Four different bed temperatures of 80, 60, 50, and 25 °C were taken into account. According to the monitoring of the temperature during the process, upon deposition of the first layer of the materials (PA6 and CF-PA6), the monitored temperature values ascended to the values of crystallization temperature .The dimensional stability/accuracy of the printed specimens was evaluated by means of optical microscopy. Based on the performed visual comparisons, the manufactured parts under a bed temperature of 80 °C had less dimensional accuracy due to the observed inconsistency and dissonance concerning the thickness of the deposited PA6 layers . The staInvestigations regarding the impact of some important FFF processing parameters of the polymer-based composites on the inter-layer adhesion (bonding) of the deposited filaments were performed. It was proven that temperature profile evaluation of the printed layers has a significant effect on the bonding of adjacent filaments. Failure stress/strain can be the indicators to determine the mechanical properties of FFF-manufactured products. In this section, using optimized processing parameters , the mechanical properties of the fabricated composite material under monotonic and fatigue loadings were analyzed . As partUsing the Mark Two printer, the tensile specimens were printed into three fill patterns: triangular, rectangular, and hexagonal. Comparisons were made based on the tensile strength of the manufactured samples made of CF-PA6. Therefore, the obtained tensile strength values of the samples concerning the stated fill pattern were compared to each other as well as to the solid fill pattern (whose infill percentage was 100%) .The obtained tensile strength of the manufactured CF-PA6 specimens in the triangular, rectangular, hexagonal fill patterns, and the solid fill were 18.79 ± 1.19, 19.84 ± 1.66, 19.99 ± 1.32, and 30.31 ± 5.5 MPa, respectively . The tenCF-PA6 filaments with continuous glass fiber were utilized to manufacture the specimens in the stated fill patterns: triangular, rectangular, and hexagonal, as well as in the solid fill pattern (whose fill percentage was 100%). Using CF-PA6 reinforced with continuous glass fiber, the obtained tensile strengths concerning the hexagonal, rectangular, triangular fill patterns, and the solid fill specimens were 60.04 ± 4.16, 66.7 ± 2.63, 69.6 ± 3.05, and 78.35 ± 0.87 MPa, respectively. Therefore, the highest tensile strength is related to the solid fill samples. In addition, by shifting the fill pattern of the printing process of the samples, which were made of CF-PA6 reinforced with continuous glass fibers in the hexagonal infill pattern (the weakest obtained strength) and the solid fill, the tensile strength improved almost 30.5% . The highest tensile strength was related to the solid fill samples .Regarding the effect of the density of the continuous reinforcement layers, a rectangular sample with thickness of 3.5 mm was designed as the tensile test specimens . The conAs was found, by increasing the quantities of the continuous glass fiber layers, the tensile strengths of the manufactured samples were increased. The tensile strengths significantly increased by enhancing the quantity of the continuous glass fiber layers from two to eight. However, the tensile strength increments of the manufactured samples from eight to ten reinforced layers were less than the related increases in strength provided by enhancing the quantities of the continuous glass fiber layers from two to four. These obtained results are correlated to the weaker adhesion strength between the glass fiber reinforcement layers and the polymer matrix layers in comparison with the polymer-to-polymer layers. By enhancing quantity of the reinforcement layers from two to eight, the tensile strengths of the samples were increased because of the existence of the continuous reinforcement, which resists against the applied tensile stress. The nature of the weaker interface strength of the glass fiber reinforcement layers and the polymer matrix layers in the manufactured samples, which consists of more than eight reinforcement layers, dominated the impact of the existing reinforcement. Thus, we can consider that in our manufactured samples, the fabricated specimens with eight layers showed the optimal density in the continuous reinforcement.According to the recently stated results, concerning the performed study on the effect of the density of the continuous reinforcement, it was concluded that there is an optimal quantity layer for the continuous reinforcement layers in the manufactured reinforced polymer composites via the FFF process. In the case of our rectangular tensile test sample, eight layers of the continuous glass fiber layers prepared an optimal tensile behavior.Therefore, the continuous glass fiber-reinforced PA6 composite samples, reinforced by eight layers, were considered for studying the tensile behavior of the manufactured samples with different continuous reinforcement directions. As for studying the effect of the continuous reinforcement direction, five direction values of 0°, 30°, 45°, 60°, and 90° were considered for continuous glass fibers. The tensile strength values of the related manufactured samples with continuous directions of 0°, 30°, 45°, 60°, and 90° were 226.34 ± 5.4, 46.8 ± 2.05, 37 ± 1.1, 36.81 ± 1.1, and 36.6 ± 0.79 MPa, respectively .As is observed, the tensile strength of the manufactured samples with continuous reinforcement in the 0° direction had the highest strength in comparison with the other fabricated samples under the different reinforcement directions. As the direction of the continuous reinforcement increased from 0° to 90°, the tensile strength decreased. A significant drop in tensile strength was observed by the change in continuous reinforcement direction from 0° to 30°. The tensile strengths of the manufactured samples with 45°, 60°, and 90° as the reinforcement directions were close to each other. In addition, the tensile strength of the fabricated samples with a continuous reinforcement direction of 90° was close to the strength of the unreinforced samples (33.65 ± 2.8 MPa). Thus, it was found that the manufactured samples with a reinforcement direction of 0° had the highest tensile strength, while as the direction of reinforcement decreased, the tensile strength decreased as well. In the case of the samples with a reinforcement direction of 0°, the glass fibers resisted against the tensile stress well because it was in the direction of the applied strength. The fabricated samples with the reinforcement direction of 90° had the weakest strength because of the perpendicular directions of the applied stress and reinforcement. Thus, in the case of the samples with the glass fiber direction of 90°, the impact of the existence of the reinforcement on the tensile strength was not so manifest and sensible. The close tensile strength value of the specimens with a reinforcement direction of 90° to the unreinforced specimens can be evidence of this phenomenon. The macroscopic fracture observations exhibited the same fracture direction as the continuous reinforcement direction .Wöhler curves for CF-PA6 and CF-PA6 reinforced with continuous glass fiber samples are shown in Based on their plots, it is evident that for high loading amplitudes, the dynamic modulus decreases rapidly in a linear regime of the logarithmic curve until the specimen fails. The dynamic modulus of applied low amplitudes deviates into three decreasing regimes–a swift one during the initial cycles (I), a gradual one (II), and a drastic decline (III) just before fracture. In addition, it was found that there is an upsurge in damage kinetics in the case of CF-PA6 reinforced with continuous glass fibers specimens compared to CF-PA6 specimens. This can be due to the existence of the weak polymer–continuous reinforcement interfaces in these specimens .Variation in the amplitude of the applied loading results in self-heating for CF-PA6 and continuous glass fiber-reinforced CF-PA6 during the fatigue test. This phenomenon affects the viscosity of the polymer as a function of temperature. It was observed that the induced temperature of CF-PA6 increased up to about 60 °C, which is in the glass transition zone. Consequently, the stiffness of the matrix agent (polymer) decreased based on the obtained loss factor evolution versus the temperature curve from the performed DMTA test, as shown in Therefore, in addition to the development of damage, the fatigue failure of the CF-PA6 specimens was caused by the viscous behavior evolution of the polymer agent and the consequent brittle–ductile transition. It is vital to consider the induced temperature during the fatigue. The maximum induced temperature during the fatigue cycles is a crucial point for considering prognosticating the fatigue behavior of the printed polymer and the continuous fiber-reinforced polymer composite materials. The maximum induced temperatures during the fatigue test of the 3D-printed CF-PA6 and reinforced CF-PA6 with continuous glass fiber samples were measured, which are illustrated in The primary purpose of this research was to study the rheological characteristics of the materials during and after the FFF process. As a result of this study, we can gain a better understanding of the behaviors of polymers and polymer composite materials, as well as how reinforcements function during and after the FFF process. The obtained results will enable us to improve the FFF-processed parts. The fundamental problem of the produced parts by the FFF process is the presence of the porosities/voids inside the manufactured parts and the subsequent poor adhesion between the deposited layers. In this study, influences of the various processing parameters on the material behaviors during and after the FFF process were investigated. The physicochemical, thermal, and mechanical assessments exhibited that the crystallinity degree was influenced by the processing parameters modification. The crystallinity degree modifications could affect material diffusion during the cooling stage and the subsequent bonding of the deposited adjacent filaments. The strength of this bonding directly affects the mechanical behaviors of the manufactured specimens. Therefore, a precise and local measurement of the temperature on the scale of the diameter of the filaments during the FFF process was performed. In addition, by comparing the obtained temperature evolution curves from PA6 and CF-PA6 specimens, it was found that the existence of the chopped carbon fibers could affect the temperature evolution curves during the FFF process.-The FFF-processed CF-PA6 specimens with solid fill patterns had superior mechanical properties and stiffness under tension.-Based on the study of the effect of the continuous glass fiber reinforcements on the fatigue behavior of the manufactured specimens, a reduction in fatigue life was observed in the manufactured CF-PA6 and continuous glass fiber-reinforced CF-PA6 specimens due to an increase in the induced temperature. By considering this induced temperature, known as self-heating, we can determine how a polymer behaves viscously, especially at the glass transition zone. It was observed that CF-PA6 reinforced with continuous glass fiber specimens exhibited a lower level of self-heating compared to CF-PA6 specimens without continuous glass fiber reinforcement during the fatigue test. Based on the microscopic analysis of the damage mechanisms during the performed fatigue tests, the fracture surface of the continuous glass fiber-reinforced CF-PA6 specimens was characterized by continuous glass fiber pull-out and subsequent debonding of the deposited layers.-As for the study of the density of continuous glass fiber in the FFF process of polymer-based composites, by increasing the density of the reinforcement layers, the strengths of the manufactured specimens were increased until a specific density value, and then, they became stable.-As for studying the effect of the direction of the continuous reinforcement in the FFF-processed polymer composite materials, the five directions of 0°, 30°, 45°, 60°, and 90° were investigated. The highest strength was related to the manufactured parts, in which the reinforcements were in the applied stress direction (0°). As the direction of the utilized reinforcement deviated more from the applied stress direction, the strength of the manufactured parts decreased. The fracture of the tensile-tested samples occurred in the direction of the deposited continuous reinforcement.The obtained results from the three-point bending fatigue and the tensile behavior of the manufactured composite materials reveal:In the next study, the effect of other continuous reinforcements such as carbon fiber and Kevlar fiber on the tensile and the fatigue behaviors of the printed specimens will be studied. Therefore, their effects will be compared with each other.


The targeted agents capmatinib and tepotinib provide a new treatment for patients with non-small cell lung cancer (NSCLC) with MET exon 14 skipping mutation (METex14). However, drug-induced pneumonitis is an uncommon but threatening adverse effect found in patients treated with both capmatinib and tepotinib. The safety of treating a patient with a MET inhibitor after drug-induced pneumonitis by another MET inhibitor remains unclear. Here, we present a case of a patient with NSCLC harboring a METex14 who was treated with a standard dose of tepotinib after advanced capmatinib-induced pneumonitis and did not present pneumonitis relapse. Tepotinib may be a safe option when medical professionals consider switching MET inhibitors after patients experience pneumonitis. Clinical Practice PointsDrug-induced pneumonitis is an uncommon but life-threatening adverse effect of target therapies. The safety of using another MET inhibitor after discontinuing the first remains unknown and such patients have rarely been reported.Given the lack of targeted therapies for MET inhibitors, rechallenge with tepotinib after capmatinib-induced pneumonitis could be an option for patients with NSCLC harboring METex14. MET exon 14 skipping mutation (METex14) occurs in approximately 3% of patients with non-small cell lung cancer (NSCLC) . TargeteDrug-induced pneumonitis, also known as interstitial lung disease (ILD), is an uncommon but life-threatening adverse effect of target therapies . While tHere, we present a case of a patient with NSCLC with METex14 who experienced pneumonitis during capmatinib treatment and was safely switched to tepotinib without pneumonitis relapse.A 69-year-old Taiwanese man without underlying lung disease and smoking history complained of right arm pain for one month. The patient was subsequently diagnosed with metastatic carcinoma based on a specimen from the operation that the patient received for treating a pathological fracture of the right humerus. NSCLC  with bone metastases was confirmed . Next-geUnfortunately, progressive exertional dyspnea and dry cough were reported after one month of capmatinib treatment. Physical examination showed bilateral basal crackles on lung auscultation. The patient’s oxygen saturation was 92% in room air, and chest radiography (CXR) and computed tomography (CT) showed increased interstitial infiltration, which was compatible with drug-induced pneumonitis/ILD . PolymerThereafter, the patient participated in a clinical trial of a novel MET inhibitor for two weeks. The patient’s pneumonitis did not relapse, but he was withdrawn from the study because of clinical progression (the metastatic tumor over the right arm progressed rapidly during treatment). No response to a novel MET inhibitor may result from the occurrence of MET resistance or no activity of this compound. Palliative chemotherapy with pemetrexed and cisplatin was immediately administered. Tepotinib, another MET inhibitor, has been reported to be active in treatment-naïve or previously treated patients with METex14 mutation. As this patient may have still benefited from MET inhibitors, tepotinib with a standard dose (one tablet of tepotinib is 225 mg and 2 tables daily is recommended dose) was also prescribed concurrently for one month and discontinued because of clinical progression (right arm tumor). There was no pneumonitis relapse within one month of tepotinib treatment .Currently, only two MET-selective inhibitors (capmatinib and tepotinib) have been approved by the U.S. Food and Drug Administration. Owing to the limited treatment options and considerable efficacy of MET inhibitors in clinical studies ,4, it isIn the current report, pneumonitis was found to occur as an uncommon adverse effect of MET inhibitor use for treating NSCLC with uncommon mutations. Therefore, the experience presented in this case report can help inform clinicians about the safety of switching MET inhibitors due to pneumonitis. Whether the dose intensity of MET inhibitors correlates with lung toxicity remains unknown. According to a previous case report, low-dose tepotinib is safe for patients with a history of grade 2 capmatinib-induced pneumonitis . HoweverThis case report had several limitations. First, no bronchoscopy study was performed to exclude other causes of pneumonitis. Furthermore, capmatinib-induced pneumonitis was diagnosed based on the rapid improvement of clinical symptoms and CXR after capmatinib discontinuation and steroid therapy initiation. Second, a longer duration of tepotinib therapy should be applied to determine whether pneumonitis relapse occurs after the long-term use of teoptinib. In this case, the patient presented pneumonitis within one month of treatment with capmatinib; therefore, we presume that pneumonitis relapse would occur sooner if an alternative MET inhibitor targeting the same molecule induces the same adverse event . NotablyWe present the case of a patient with NSCLC with a METex14 alteration who experienced capmatinib-induced pneumonitis and safely continued treatment by switching to alternative MET inhibitors, including one from a clinical trial and tepotinib. Given the lack of targeted therapies for MET inhibitors, rechallenge with tepotinib after capmatinib-induced pneumonitis could be an option for patients with NSCLC harboring METex14. Our experience suggests that tepotinib may be a safe therapeutic option for patients who have recovered from capmatinib-induced pneumonitis under close clinical surveillance.


Continuously reinforced concrete pavement (CRCP) is a representative type of concrete pavement constructed with continuous steel bars without intermediate transverse expansions. With reference to pavement conditions, CRCP is an exceptional type of concrete pavement according to the Highway Pavement Condition Index (HPCI) and International Roughness Index (IRI). The two main design methods for CRCP are AASHTO 86/93 and the Mechanistic–Empirical Pavement Design Guide (MEPDG). Because of limitations of the AASHTO 86/93 design method, the MEPDG method is more reliable. While incorporating the interactions among geometrics, pavement structure layers, material properties, subgrade, traffic, and environmental conditions, and the prediction values according to the MEPDG method, it matched the measured results of crack spacing and crack width. The MEPDG punchout, crack width and spacing, and load transfer efficiency (LTE) models were evaluated, and results were compared with the test sections in three European countries consisting of different construction details, which were investigated and recorded between 2019 and 2021. In this sense, a calculation tool was developed to consider the influence of different parameters in design process. In addition, sensitivity analyses were executed for the development of punchout, considering various input parameters. The track surveying and the evaluation of the results indicated that the design process can be improved with consideration of some criteria such as crack formation time or adjustment of the correlation between crack width and crack spacing. Due to the very important function of erosion and resulting pumping in the deterioration of CRCP, it is advisable to include the influence of the base layer and the influence of different shoulder type and heavy traffic volume or effect of deflection in the calculations. Continuously reinforced concrete pavement (CRCP) is a characteristic type of concrete pavement constructed with continuous steel bars without intermediate transverse expansions. In terms of pavement conditions, CRCP is a superior type of concrete pavement according to the Highway Pavement Condition Index (HPCI) and International Roughness Index (IRI). The steel reinforcement manages the occurrence of random transverse cracks resulting from volumetric changes due to cement–water hydration and holds the cracks tightly together such that the concrete slab behaves like a continuous system to transfer loads across the crack interface ,2,3. TheThe reason for these discrepancies between predicted MEPDG values and measured values is the deficiencies in the design method, which have important functions in the design of CRCP and are not considered in the procedure of MEPDG .The objective of this publication  is to idPCC,i, modulus of rupture MRPCC,i, tensile strength ft,i, and compressive strength fc,i) are calculated monthly, whereas traffic load and ambient temperature are calculated up to hourly. The drying shrinkage has a significant influence on the crack characteristics that develop at an early age and depend on the humidity in the concrete. Another important component for the CRCP calculations is the drop in PCC temperature at the depth of steel ΔTmζ,. The modulus of elasticity of the subbase (SBE) and the modulus of subgrade reaction (k) are also important in the MEPDG design method. The last time-dependent parameter to be determined is the radius of relative stiffness il This value is used to describe the relative stiffness of the concrete slab compared to its foundation. The crack spacing is dependent on CRCP age, traffic load, and fatigue, and its calculation in the MEPDG includes the tensile strength of the concrete after 28 days ft (psi), the Bradbury’s curling/warping stress coefficient C, Westergaard’s nominal stress factor σ0 (psi), the depth to steel layer ζ (in), the slab thickness PCCh (in), the coefficient of base friction f, the peak bond stress mU (psi), the steel ratio bP, the coefficient of first bond stress c1, and the reinforcing steel bar diameter bd (in). At an early age, temperature and moisture variations induce volume changes in the concrete, which are arrested by the reinforcement and the base friction, leading to the development of stresses and occurrence of transverse cracks, because the tensile stress exceeds the concrete’s strength , PCCα is the PCC coefficient of thermal expansion [°F−1], Δζ,mT is the drop in PCC temperature from zero-stress temperature at the depth of the steel at the time of crack width prediction [°F], c2 is the coefficient of second bond stress, depending on concrete strength and crack spacing by 0 °C at the level of reinforcement and zero temperature differential, EPCC is the concrete elastic modulus, and σf is the maximum longitudinal tensile stress in the concrete at the level of reinforcement. The above equation provides the crack width according to total strain (from shrinkage and temperature drop) at the level of steel and the restrained longitudinal deformation caused by stresses in the concrete.The average crack width coming from MEPDG is calculated as follows :(1)cwi=CiPO is the total predicted number of punchouts at the end of the time increment, POFD is the accumulated fatigue damage, and A, α, and β are the calibration constants for the punchout function. The maximum top of concrete transverse tensile stress is calculated depending on the traffic level and axle load distributions for each time increment. Damage is calculated on the basis of the stress level and traffic load applications. This damage is accumulated over time, and the possibility of punchout is calculated.For CRCP, major structural distresses considered for design are punchouts, which are a result of water infiltration due to wider crack opening in combination with erosion of the base layer due to pumping effects, resulting in a CRCP pavement which is no longer completely supported; therefore, additional stresses appear in the concrete with closer transverse cracks and longitudinal cracks occurring ,21,22. TBetween 2019 and 2021, sections in Germany, Belgium, and Poland were investigated in the field, along with the help of extracted drill cores in the laboratory . A5 BrucThe time of crack assessment and evaluation of the crack spacing are presented in Equation 1) assumes that the cracks occur directly after the concrete is placed and no more cracks occur over the course of time, although the formation of cracks is completed after 2–4 winter periods, depending on the climatic conditions , C1 and C2 are constant factors, PO is the number of punchouts [1/mi], and SF is the site factor.The majority of damage to CRCP in the form of punchouts is caused by quality problems in the materials or substandard construction practice near the joint and the repair joint, and this issue is not involved in the design method, although, in the calculations of the IRI, the condition of the pavement in the form of initial as-constructed IRI, as well as the site factor, is included in the algorithm:As indicated in Conversely, the LTE is only moderately sensitive to reinforcement ratio and base type, while it is insensitive to the coefficient of thermal expansion (CTE), steel depth, and base friction. The value of LTE depends on the type of base layer, reinforcement ratio, and crack stiffness. LTE is not a reliable indicator for predicting the punchout potential during regular construction regardless of a slab thickness and crack spacing higher than 90%, and with sufficient longitudinal reinforcement. Additionally, LTE values for a crack where a punchout process was performed remained at a relatively high level before and right after the development of the punchout .c,i).The calculation for LTE adopted in the MEPDG is influenced by constant variables which are contributed by the base layer; the only parameter that changes over time and affects the LTE calculations is the variation in transverse crack stiffness  in an interval of 20 m in the longitudinal direction . The anaThe calculations of LTE according to the MEPDG for the A5 near Darmstadt also do not agree with FWD measurements in a timeframe of 15 years after construction of the CRCP. The evaluation of FWD measurements showed a reduction in bearing capacity on sections III, IV, and V of A5 near Darmstadt, although, according to the MEPDG, section IV consisting of cement-treated base should provide satisfactory LTE up to the 23rd year after installation. In the case of section II consisting of asphalt base course, the bearing capacity increased after 15 years, although, according to calculations, the LTE should fail after 7 years .iRE is the monthly rate of base erosion from the slab edge [in/month], P200, is the percentage subgrade passing the no. 200 sieve, EROD is the erodibility index depending on the material from 1 to 5, and PRECIP is the mean annual precipitation [in]. The erosion based on Equation (4) is completely dependent on properties of the base layer, and the influence of different shoulder types and heavy traffic volumes is not included. The effect of deflection in the calculations is not considered, although the deflection measurements on the investigated sections proved the erosions.The base layer is critical in terms of transmitting the stresses and maintaining the stability of the structure, and it influences the behavior of the CRCP on different aspects such as friction between slab and base layer or erosion. On one hand, a high friction will lead to higher stresses and a homogeneous crack spacing. On the other hand, a total bond between the pavement and the cement-treated base layer should be avoided as they will act as one, with a consequently reduced reinforcement rate. The placement of a geotextile between the CRCP and cement-treated base layer will accordingly act as a stress relief layer and allow both layers to deform separately. Although the implementation of geotextile on CTB resulted in a uniform crack pattern and erosion traces were not observed, this was not considered in the MEPDG design. Additionally, requirements for base courses are usually not answered within the scope of computational dimensioning, but enter into the procedures obliquely via input parameters such as the modulus of subgrade reaction.Due to significant discrepancies between predicted MEPDG values  of the re-evaluated sections and measured values on the sites, in this study, the limitations of the MEPDG design process were identified, and the punchout and LTE predictions, as well as crack width models, for CRCP characterized by MEPDG design were evaluated. The formation of transverse cracks is completed during the shrinkage period, and crack formation is not finished directly after the placement of concrete. The calculation of the crack width can be improved by the consideration of the concrete shrinkage as a function of time. In this way, the influence of crack formation during the shrinkage process or after completion of the shrinkage period can be incorporated. Moreover, in the crack width calculations, the concrete area contributing to the crack width can be limited. The linear correlation between the crack spacing and the crack width needs to be adjusted. After the evaluation of the FWD values on the test tracks, it is recommended to consider the initial as-constructed and measurable values such as FWD in the calculations for the prediction of the punchout or LTE. Due to the very important function of erosion and resulting pumping in the deterioration of CRCP, it is advisable to include the influence of the base layer and the influence of different shoulder type and heavy traffic volume or the effect of deflection in the calculations. LTE is quite high during regular construction before and shortly after the development of punchouts; with sufficient longitudinal reinforcement, it is not a reliable indicator for predicting the punchout potential. The effect of reinforcement cover and the month of construction on the development of punchouts is not directly included in the design calculations, although this impact has already been confirmed. An important parameter contributing to the development of punchouts is the presence of horizontal cracks in the reinforcement layer, which, despite confirmation of their impact, have not yet been imported into the calculations.


The shear strength provisions used by some international codes, such as the American Concrete Institute (ACI-2019), the Eurocode 2 (EC-2), and the Egyptian Code (ECP 207), were examined when applied to HSC beams with and without NS. The most important factors to consider were the effect of using NS on the shear span to effective depth (a/d) ratio and the shear strength of the beams with and without stirrups. The experimental results were validated using a nonlinear finite element analysis using the computer program ABAQUS. The experimental results showed that increasing the NS ratio reduced the number of cracks, and increased the cracks spacing, as well as reducing crack width. In specimens without stirrups, these effects were more obvious. A rise in the (a/d) ratio increased the number of cracks along the beam length, notably in the mid-span region. For specimens without stirrups and with an (a/d) of 1.5, raising NS from 0% to 1%, 2%, and 3% increased the ultimate load by 13%, 30%, and 39%, respectively, whereas for specimens with an (a/d) of 2.5, the ultimate load increased with approximately the same increase as that in beams with an (a/d) of 1.5 due to using NS. Additionally, the addition of NS to concrete boosted the contribution of the concrete to the shear strength, as shown by the results of beams without stirrups. For specimens with stirrups and an (a/d) of 1.5, raising NS from 0% to 1%, 2%, and 3% increased the ultimate load by 8%, 21%, and 30%, respectively. Additionally, for specimens with stirrups and an (a/d) of 2.5, the ultimate load increased with approximately the same increase as that in beams with stirrups and an (a/d) of 1.5 due to using NS. The test findings indicate that the shear strength calculated using the equations of the ACI 318-19 is more conservative than EC-2 and ECP 207 for NS concrete beams. The finite element program ABAQUS may be successfully used to predict the shear strength of NS concrete beams.In this study, the shear strength of sixteen full-scale over-reinforced concrete beams with and without nano silica (NS), constructed from high-strength concrete (HSC), was investigated both experimentally and analytically. Nano silica was used as a partial replacement for Portland cement. According to the NS ratio, the tested beams were divided into four groups: 0%, 1%, 2%, and 3%. Shear span to effective depth (a/d) ratios of 1.5 and 2.5 were used in each group, and two different stirrups ratios (ρ Nowadays, concrete is considered a widespread worldwide construction material. However, concrete is a brittle material and is distinguished by a very low tensile strength. The use of nano silica (NS) for supplementing cementing materials is a new technology that has been applied to concrete manufacturing to enhance the mechanical properties of hardened concrete . In geneIn this research, the shear strength of sixteen full-scale over-reinforced concrete beams with and without nano silica (NS), constructed from high-strength concrete (HSC), is investigated both experimentally and analytically. When applied to HSC beams with and without NS, the shear strength provisions used by several international codes, such as the American Concrete Institute ACI-2019) , the Eur , the Eucu based on three cubes (150 × 150 mm) were determined, as illustrated in CEM I 52.50N ordinary Portland cement was utilized. The coarse aggregate was crushed limestone with a size of 8 mm. Natural sand with a fineness modulus of 2.80 was used as the fine aggregate. Nano silica (NS) is an amorphous silica powder that is exceptionally pure and has a large surface area, outstanding adsorption properties, and excellent dispersion. NS is a fine material with a diameter of 50 mm; this makes it 50 times smaller than cement. As a result, adding NS to a lower level of cement replacement improves the mechanical and durability properties of concrete by accelerating the hydration and pozzolanic reaction at an early stage. v) were employed at 0% and 0.38%. The stirrups were constructed from 6 mm-diameter bars with 200 mm and 100 mm spacing. For all the tested beams, the main bottom reinforcement was over-reinforced with four bars with a diameter of 18 mm to avoid flexural failure and two bars with a diameter of 12 mm as a top reinforcement. The flexural reinforcement ratio (ρL) for all tested beams was equal to 3.53%. The design yield stress was 410 MPa for longitudinal bars and 243 MPa for stirrups. Longitudinal bars had a design yield stress (yf) of 410 MPa, whereas stirrups had yield stress (ysf) of 243 MPa. All specimens were cast and cured with purified water for 28 days before testing. Sixteen simply supported HSC beams with different percentages of NS were experimentally examined up to failure under two-point symmetric loading. All the tested beams had concrete sections of 150 × 250 mm with lengths of 1600 mm. The reinforcement details and concrete dimensions of the tested beams are illustrated in A conventional universal machine was used to test all beams. A load cell with a capacity of 300 kN and an accuracy of 0.1 kN was used to measure the loads. The beams were tested under two-point loads with two shear span–depth ratios of 1.50 and 2.5, respectively. cr), ultimate load (Pu), and mid-span deflection (∆u).For all tested beams, The relations between loads and longitudinal steel strains were approximately linear, since all beams were over-reinforced and the lower longitudinal steel for all specimens was not yielded. Furthermore, the results demonstrated that increasing NS reduced tensile steel strain and concrete compressive strain at various load levels. The load–tensile stirrups strain correlations in the vertical legs of the stirrups passing through the failure plane (about in the middle of the shear span) are shown in uv) in a concrete beam reinforced with stirrups was resisted by the concrete (cv) and the stirrups (sv).λ is equal to one for normal concrete; cf′ is the cylinder concrete compressive strength in (MPa); bw is the concrete beam width in mm; d is the effective beam depth in mm, within the spacing, s; Av denotes the area of shear reinforcement; ysf is the yield strength of shear reinforcement (MPa).The shear critical section was at a distance of d from the column face . The totcv) was calculated as follows:ckf denotes the concrete’s characteristic compressive cylinder strength after 28 days, and cdf is the concrete compression force design value in the longitudinal beam axis direction. Shear strength resisted by stirrups (sv) should be taken as the lesser of sv,1 or sv,2:sw is the cross-sectional area of the shear reinforcement. ywdf is the shear reinforcement’s design yield strength.The critical section for shear was located at a distance of 2d from the column face . The desuv) is calculated as ACI 318-19, except the shear strength resisted by the concrete (cv), which is calculated as follows:cuf is the cube concrete compressive strength and γc is the safety factor for concrete and taken as equal to 1.50.The shear critical section is at a distance of d/2 from the column face . The totuexpv) of the tested beams were compared with that estimated from ACI 318-19 (uACIv), EC-2 (uECPv), and ECP 207 (uECPv) codes, as illustrated in uexpv/uACIv), EC-2 (uexpv/uECv), and ECP 207 (uexpv/uECPv) were 1.64, 1.41, and 1.36, respectively.The envelopes of the experimental shear failure stress  ratio led to an increased number of cracks.Raising NS from 0% to 1%, 2%, and 3% increased the ultimate load by 13%, 30%, and 39%, respectively, for beams without stirrups and with (a/d) of 1.5, whereas for specimens without stirrups and with (a/d) of 2.5, the ultimate load increased by 15%, 31%, and 36%, respectively.Raising NS from 0% to 1%, 2%, and 3% increased the ultimate load by 8%, 21%, and 30%, respectively, for beams with stirrups and with (a/d) of 1.5, whereas for beams with stirrups and with (a/d) of 2.5, the ultimate load increased by 7%, 22%, and 31%, respectively.Raising NS from 0% to 1%, 2%, and 3% increased the mid-span deflection by about 6–18%.The ACI 318-19 is more conservative than EC-2 and ECP 207 in calculating the shear strength of NS high-strength concrete beams.The finite element program ABAQUS can be used effectively to predict the shear strength of nano silica-reinforced high-strength concrete beams.


Along with the increasing global burden of diabetes, diabetic foot infections (DFI) and diabetic foot osteomyelitis (DFO) remain major challenges for patients and society. Despite progress in the development of prominent international guidelines, the optimal medical treatment for DFI and DFO remains unclear as to whether local antibiotics, that is, topical agents and local delivery systems, should be used alone or concomitant to conventional systemic antibiotics. To better inform clinicians in this evolving field, we performed a narrative review and summarized key relevant observational studies and clinical trials of non-prophylactic local antibiotics for the treatment of DFI and DFO, both alone and in combination with systemic antibiotics. We searched PubMed for studies published between January 2000 and October 2022, identified 388 potentially eligible records, and included 19 studies. Our findings highlight that evidence for adding local antibiotic delivery systems to standard DFO treatment remains limited. Furthermore, we found that so far, local antibiotic interventions have mainly targeted forefoot DFO, although there is marked variation in the design of the included studies. Suggestive evidence emerging from observational studies underscores that the addition of local agents to conventional systemic antibiotics might help to shorten the clinical healing time and overall recovery rates in infected diabetic foot ulcers, although the effectiveness of local antibiotics as a standalone approach remains overlooked. In conclusion, despite the heterogeneous body of evidence, the possibility that the addition of local antibiotics to conventional systemic treatment may improve outcomes in DFI and DFO cannot be ruled out. Antibiotic stewardship principles call for further research to elucidate the potential benefits of local antibiotics alone and in combination with conventional systemic antibiotics for the treatment of DFI and DFO.  Diabetic foot infections (DFI), including diabetic foot osteomyelitis (DFO), represent a societal challenge associated with substantial morbidity, prolonged hospitalizations, prospective amputations, and higher overall healthcare costs . DiabeteIt is worth mentioning that in the local treatment of DFI and DFO, aminoglycosides and vancomycin are the most studied antibiotics, having both broad-spectrum activity and being able to be incorporated into any delivery vehicle. The biological properties of these two antibiotics also allow them to remain thermostable during the exothermic reaction that occurs during the polymerization of the cement in local applications, ultimately delivering a significantly higher concentration locally than via the blood system . For the purposes of this review, we consistently apply the term topical to refer to the local administration of antibiotics through an infected diabetic foot ulcer (DFU). Topical antibiotic interventions have previously been studied in the assessment of the bioburden and biofilm of diabetic foot wounds . For exaInterestingly, in vitro studies have shown that the kinetics of antibiotic release from gentamicin collagen sponge matrices can reach 95% in as little as 1.5 h ,14. A pl®, Simplex®, Eurofix®, Palacos®, Copal®, and Refobacin®. As with all non-absorbable materials, the main disadvantage of PMMA is its surgical removal, resulting in additional intervention following the release of all drugs. Alternatively, there are hybrids of biodegradable carrier systems that take advantage of different properties to improve local antibiotic release. They have different presentations, such as Cerament G/V® —an acrylic used extensively in orthopedic surgery for chronic osteomyelitis and implant-related infections ,19. Some®  for the treatment of bacterial DFI and DFO, both alone and concomitant to systemic antibiotics. Because some systematic reviews have been published without any definitive conclusions regarding the use of local antibiotics in DFI and DFO , we beliWe identified 388 potentially eligible studies following our search strategy. All studies were screened and assessed for eligibility by the lead author (LSB), and only those that met the eligibility criteria were included in the analysis. 19 studies were finally included in our review: 5 randomized clinical trials (RCTs) related to topical antibiotics, and 14 studies (1 RCT and 13 retrospective studies) related to local delivery antibiotics in DFI and DFO. In general, retrospective studies exploring local antibiotic delivery had methodological shortcomings and included highly heterogeneous study populations. Hence, most studies lacked comparability with each other and were not suitable for formal quantitative pooling via meta-analysis . A remarOur pragmatic narrative synthesis stressed that chronic open wounds of the diabetic foot are often complicated by infection and, together with advanced ischemia, can result in amputations . Some ofp < 0.05). Varga et al.  performeOne of the included multicenter RCTs, consisting of two consecutive double-blind trials, compared topical versus systemic antimicrobial therapy for the treatment of mildly infected DFU . A totalAnother relevant pilot RCT for mild DFI with ulcers (n = 22), performed within the Swiss context, compared local gentamicin sponges to standard wound care without systemic antibiotics. At the primary endpoint, 20 patients (91%) were considered to have reached a clinical cure for infection, and 2 (9%) were deemed to have achieved significant improvement. In terms of microbiological outcomes, only 12 patients (56%) achieved microbiological eradication of all pathogens. This study did not find any significant between-group differences in either clinical or microbiological outcomes, yet it provided reassurance about the tolerability and feasibility of adding gentamicin-sponges to standard wound care .A pragmatic RCT for moderate to severe DFI (n = 56) compared a topical gentamicin collagen sponge plus a systemic antibiotic intervention versus systemic antibiotics alone. No significant between-group differences were observed in the main outcomes. However, this study is not directly comparable to the other included studies in our narrative review due to the design of the combined local and systemic antibiotic intervention and the underlying research question of interest . A similar RCT based in Switzerland compared gentamicin-collagen sponges in combination with systemic therapy to standard wound care with systemic therapy in 88 diabetic patients with moderate to severe DFIs. Although the gentamicin sponge intervention was well tolerated, it did not show superiority to standard wound care. Nonetheless, a suggestive trend toward more rapid and pronounced ulcer healing was observed in the gentamicin-sponge during the second half of the study (from 3 to 5 weeks) .The reviewed body of evidence regarding local delivery systems for forefoot DFO is limited to one RCT. Forefoot DFOs have an inherent higher frequency of therapeutic surgical amputations and are always at risk of postoperative skin breakdown and subsequent major amputations ,28. In cNiazi et al.  publisheIn contrast, one small RCT of 36 participants with diabetic foot ulcers showed that coadjuvant local antibiotics helped shorten healing times . SimilarConservative treatment in midfoot-hindfoot DFO is less frequently investigated than forefoot DFO . That saOur narrative review underscores that there is very limited evidence favoring the addition of local antibiotic delivery systems to standard DFI and DFO treatments. Experimental and observational studies of local antibiotic interventions published to this date have more frequently targeted forefoot DFO, although the marked heterogeneity of the study populations in the included studies precludes the drawing of robust general statements. Importantly, some of the included studies in our narrative synthesis reveal that the addition of local agents to conventional systemic antibiotics might help in key outcomes of a diabetic foot, namely time to recovery and clinical cure proportion. Critically, interventions consisting of local antibiotics as a standalone approach remain neglected in DFI and DFO management.Our review identifies some key knowledge gaps in the rapidly evolving area of DFI and DFO treatment and supports further research on local antibiotics. It is worth emphasizing that, given that some of the included clinical trials were likely underpowered, the lack of statistically significant results should not be interpreted as strong evidence against the clinical usefulness of local antibiotics. Furthermore, because larger-scale and appropriately powered trials have focused on comparing the addition of topical or local intraosseous antibiotics to conventional systemic antibiotics—which tend to include professional wound care and adequate off-loading —alternatOnly recently has antibiotic management found its way into the field of academic research, motivated primarily by questions related to the ideal duration of systemic antibiotic therapy ,47,48. HRemarkably, the international Infectious Diseases Society of America (IDSA) and International Working Group of Diabetic Foot (IWGDF) guidelines have a strong position against the use of topical antimicrobials in the diabetic foot, although they also promote research in this promising field. IDSA justifies its position based on limited data on topical antimicrobials for mild diabetic foot infections with minimal cellulitis ,17, whilDFIs are associated with significant therapeutic challenges, and decisions regarding their optimal therapeutic approach are difficult. In addition to surgical debridement, professional wound care, adequate off-loading, and eventual revascularization, systemic targeted antibiotics have traditionally been required when there is an infection, especially for moderate and severe DFIs.With clinical equipoise being justifiable, future trials should strongly consider head-to-head comparisons of local antimicrobial therapy alone versus systemic therapy in certain subsets of DFI and DFO . It is plausible that, in addition to shortening the treatment of DFO, topical antibiotics could be effective in mild and ulcerated DFI. A specific and potentially informative trial in the future should contemplate assessing short-term (1 to 2 weeks) oral therapy versus topical antibiotic interventions in a study population of mild diabetic foot ulcer infections or diabetic foot ulcers with secondary signs of infection. Although clinical studies of local antibiotics for DFO are more difficult to justify, some valuable information could be revealed by targeting specific subpopulations and carefully refining research questions of interest. For example, local antibiotics trials for DFO would benefit from homogeneous eligibility criteria, well-defined anatomical locations, specific surgical approaches, and consistent primary endpoints.The main limitations of this study arise from the narrative review design and the pragmatic search strategy, which may introduce subjectivity in the selection of studies. However, we judiciously and purposively selected clinical and observational evidence to reflect the relevant body of evidence to answer our research question and provided an expert interpretative overview of the field. Within the context of our review and having made suggestions for further research, it is worth citing a promising published protocol for a multicenter trial of orthopedic infection, which will use a non-inferiority design to compare local antibiotics plus 7 or less days of systemic therapy versus local antibiotics plus 4 or more weeks of systemic therapy. Remarkably, this upcoming trial will include adults undergoing surgical treatment for orthopedic infections, including a diabetic foot, and the surgical intervention will consist of a local antibiotic-carrier implantation at the site of infection . Novel pWe performed a literature review in the PubMed database in October 2022, searching for the following terms: (diabetic foot) AND . We limited our pragmatic search strategy to one database for feasibility purposes, as we considered that this would allow us to answer the research question of interest given our resource constraints. We applied filters to restrict potentially eligible titles to be published between January 2000 and October 2022.To be included, studies needed to enroll adult patients ≥ 18 years with ongoing diabetic foot problems  and local antibiotic use. We included randomized clinical trials, observational studies, and case series with > 10 patients. In addition, we hand-searched for additional citations from the references of the retrieved articles. We excluded articles written in languages other than English. Further details on our search strategy can be found in the In preparing this narrative review, we adhered to key applicable items for narrative reviews from the PRISMA statement . We alsoTopical and local intraosseous antibiotics might improve the healing of mild DFU infections or operated DFO cases, although they have been explicitly neglected as standalone interventions. Most studies published to date have failed to show the superiority of adding these local treatments to systemic antibiotic administration. Given the potential systemic antibiotic-sparing benefits of local antibiotics as standalone treatments, further research is needed to promote antibiotic stewardship principles and advance the field of DFI and DFO management.


For the more extreme (high) conditions, the use of 40 kg/m3 of RTSF fiber was sufficient to completely eliminate surface plastic shrinkage cracks. This work can help develop more sustainable concrete structures in a wider set of environmental conditions and help mitigate the impact of climate change on concrete infrastructure.Due to climate change and population expansion, concrete structures are progressively being subjected to more extreme environments. As the environment affects plastic shrinkage directly, there is a need to understand the effect of environmental changes on plastic shrinkage cracking. This paper examines the plastic shrinkage crack development parametrically at low, mid, and high drying environmental conditions, corresponding to different environments in three Saudi cities. The effects of water-cement ratios and quantities of recycled tire steel fibers (RTSF) in concrete are also investigated. The different environmental conditions for the plastic shrinkage tests were simulated in a specially designed chamber as per ASTM C1579, 2006. A digital image processing (DIP) technique was used to monitor crack initiation and development. Through the use of the crack reduction ratio (CRR), it was found that 30 kg/m Population expansion is also pushing development in more extreme environments. CO America ,2. Many  America .Long-term concrete cracking is unavoidable, and large openings impact concrete durability ,5. In hoVolume changes in concrete before the hardening of cement-based materials are the main cause of plastic shrinkage strain and cracking ,17. VoluHot weather casting is known to increase plastic shrinkage cracking . It is wAmbient conditions influence the water evaporation rate. As air temperature increases, relative humidity sees a corresponding decrease, and thus evaporation rate increases. Higher wind speeds also increase the evaporation rate. When the evaporation rate is less than the bleeding rate, a thin layer of water covers the surface of the concrete, which helps increase the evaporation rate due to the increased exposed area .Eventual drying of the surface leads to a rise in capillary pressure converting it from a mildly compressive to a tensile pressure . When ca2/h. However, Sayahi and Hedlund [2/h, especially in hot weather conditions. For example, Almusallam et al. [2/h.According to ASTM C1579 , plastic Hedlund  found thm et al.  concludeTo examine the possibility of plastic shrinkage cracking in concrete, ASTM C1579  recommenNabil et al.  examinedw/c ratio of about 0.45, with encouraging results.Almusallam et al.  and Safiw/c (0.6) due to the high-water absorption of the recycled sand. The results showed that increased replacement ratios of recycled sand mortar resulted in increased plastic shrinkage cracking. Cohen et al. [Zhang and Xiao  investign et al.  found thn et al.  came to n et al.  examinedIn conclusion, plastic shrinkage cracking is likely to worsen with climate change, and further work is needed to understand both how the environment and mitigation measures affect its development.The increase in harsh environmental conditions created by climate change can significantly affect concrete durability, in particular, due to their impact on plastic shrinkage. Hence, there is a need to investigate such conditions and their impact on concrete with various compositions, as well as possible mitigation measures.Saudi Arabia has been selected as a case study region in this investigation as it typifies extremely hot environments, with a drier climate in the northern and central areas and a wetter climate in the southern, western, and eastern areas. Despite the significant variation in humidity, most of the countries in the Middle East share the same high air temperatures .Hasanain et al.  tested tThree cities in Saudi Arabia were chosen to evaluate the effect of concreting under severe hot and dry weather conditions. (a)RiyadhThe capital of Saudi Arabia, Riyadh city, is located in the middle of Saudi Arabia. Temperatures are high in the summer, and the relative humidity is very low in both winter and summer see b. DuringArafah et al.  investig(b)DhahranDhahran city, on the eastern side of Saudi Arabia, is one of the hottest cities in Saudi Arabia and the Middle East region, and its climate is characterized by high humidity levels in both winter and summer see c as it iw/c ratios and temperatures used in the study. A high evaporation rate was observed for the plain concrete , and this was attributed to the high heat evolution, which in turn had a negative impact on the overall shrinkage behavior.Al-Gahtani  conducte(c)HailKnown as one of the coolest cities in winter in the country, Hail is one of the northern cities in Saudi Arabia. Few studies have examined concreting in this region, but while the relative humidity in this part of the country is relatively low during the months of June to September, the combination of environmental conditions experienced in this area is likely to affect plastic shrinkage ,31,43 reDue to the different levels of relative humidity in the selected case study regions, the relative humidity was kept around 20%, which is close to the minimum over the summer for Saudi Arabia see a.w/c ratios. The addition of different dosages of recycled tire steel fiber (RTSF) is also examined as a plastic shrinkage mitigation measure. The results will help improve concrete performance in hot regions such as the Middle East and in more extreme environments created by climate change and will reduce the use of natural and virgin materials by reusing tire steel fiber to control early-age cracking in concrete structures.This paper examines the impact of various environmental conditions  on the plastic shrinkage performance of concrete made with different 3 with a w/c of 0.55 (for the reference mix) superplasticizer  was used at a dosage of 1.5 lt/m3. Two sizes of gravel (river round gravel) were used, 491 kg/m3 of 5–10 mm, and 532 kg/m3 of 10–14 mm. The fine aggregate (river-round sand) was 847 km/m3. RTSF was used at different proportions .As the likelihood of plastic shrinkage cracking increases with high cement content, a relatively high cement content was selected . The cemThe water-cement ratios that were used in this study include 0.5, 0.55, and 0.6 as Low, Mid, and High ratios, respectively.Concrete mixing and specimen casting were carried out according to ,48. The Waste disposal is becoming a critical environmental pollution issue worldwide. According to Mohajerani et al. , billionHu et al.  showed tThe advantages of using RTSF instead of MSF in concrete are not just in terms of overall performance but also in terms of economic and environmental benefits, as RTSF has lower greenhouse gas emissions and cost compared to MSF ,59,60. M3 by volume can stop plastic shrinkage cracks. That same amount is used in high, mid, and low environmental conditions and increased to 40 kg/m3 if cracks develop. RTSF E=M2The test was carried out on fresh/plastic concrete in accordance with ASTM C1579  and ACI At the end of the test (after 6 h), the doors of the chamber were left open, and the slabs were covered with a plastic sheet and left undisturbed for an additional 18 h. The cracks were then measured 24 h again after casting.The method used to measure the cracks followed the recommendations of ASTM C1579 , coupledThe low wind speed of 3 m/s and the high wind speed of 7 m/s corresponds to the average wind speeds experienced in the summer in Riyadh and Dhahran, respectively.The 28-day compressive strength of the control specimens (PC) and SFRC reinforced specimens was determined on cubes that were kept both insides (In) and outside (Out) the chamber during the plastic shrinkage tests, and the corresponding values are summarized in The results do not show any significant influence of temperature and wind speed on the compressive strength of the cubes that were kept inside the chamber see a. As expNaturally, concrete strength increases with the reduction in the water-cement ratio see b in line2/h, or the test should be rejected. It can be seen that all the environmental conditions exceeded that rate and achieved the minimum value even in the low environmental conditions. As expected, in general, the evaporation rates increase with time, as bleeding gathers momentum and concrete temperature increases with the heat of hydration and then slows down towards the end of the test as bleeding slows down with setting.The evaporation rate increases with wind speed and temperature and nearly doubles from Low to High conditions. This can be attributed to the higher bleeding induced by temperature and wind, as also reported by others ,76,77,78The evaporation rates for low, mid, and high water-cement ratio mixtures are shown in For crack measurement, a DIP system was used in which two cameras were placed above the slabs, with each camera taking images every 10 min during the test. The images were subsequently analyzed using MATLAB to measure an average crack (width). A typical photo of the crack for the plain mix is shown below :w/c ratio, respectively. The graphs also show the crack widths at 24 h measured both with DIP and a conventional manual optical device. The crack widths obtained for the two methods of measurement are almost identical, confirming that DIP works well. The mean difference between the two measurements was 0.113 mm.3 of RTSF seems to prevent cracking completely for the Low and Medium temperatures. However, though the addition of RTSF delays and helps control the crack, cracking still develops at High temperatures. For the more extreme temperature tested in this study, the increase in fiber dosage to 40 kg/m3 seems to prevent plastic cracking.Similarly, to what was observed for increasing temperature values, increasing wind speed causes earlier cracking to develop and eventually leads to larger crack widths. Interestingly, both the Low wind speed and Low-temperature environments, which are below the ASTM C1579  recommen3 prevented cracking in the specimens subjected to the Low and Medium wind speeds only, whilst 40 kg/m3 of RTSF was effective in preventing cracking at High wind speeds.Again, the use of RTSF at a dosage of 30 kg/mw/c ratios. As expected, the use of lower w/c ratios delayed the crack initiation and development. As with the previous results, the addition of 30 kg/m3 of RTSF was not sufficient to prevent cracking in the specimens with a high water-cement ratio, but 40 kg/m3 of RTSF prevented the formation of cracks completely.The effect of the three parameters examined in this study on the evaporation rate at crack initiation and time of crack initiation is shown in w/c ratio, and the final cracks are wider. The plain mid-temperature concrete cracked just before the second hour, as expected by [As expected, for all mixes, cracking starts earlier, and the evaporation rate is higher with increasing temperature, wind, and ected by .Cracks develop even at the low temperature of 28 °C but crack initiation is delayed by about an hour when compared to specimens exposed to the highest examined temperature. Though surface temperature cannot be changed easily, to reduce plastic shrinkage cracking, the concrete temperature can be cooled down by using cooler materials, including iced water.In the case of high air temperature see , it was At low wind speeds (3 m/s), the evaporation rate reduces but the plain concrete cracks only a bit later and a bit less wide. Hence, at the ranges used, wind speed seems to play a less important role than temperature.w/c ratio seems to influence the evaporation rate the least, compared to temperature and wind. However, the w/c ratio has a big influence on the time of crack initiation. This was attributed by Topçu, and Elgün [w/c ratios only accelerate the initiation of high cracking due to the earlier and higher bleed rate.The nd Elgün  to a hig3 and 40 kg/m3 of RTSF. It can be seen that the addition of 30 kg/m3 of RTSF can eliminate the occurrence of plastic shrinkage cracks at low (28 °C) and mid (36 °C) temperatures.Crack reduction ratios were determined in this study according to Equation (1). In comparison, Borg et al.  examined3 of RTSF effectively reduced these cracks by up to 45%, and 55%, respectively. The addition of RTSF at a rate of 40 kg/m3 eliminated the cracks completely, and this dosage is recommended for these high temperatures.At the highest examined temperature of 45 °C (High), the plain concrete developed two surface cracks, and the corresponding CRR rations are represented in Al-Tulaian et al.  used rec3 of RTSF eliminated cracks at low and mid-wind speeds, while 40 kg/m3 of RTSF was necessary to eliminate plastic shrinkage cracking at high wind speeds.The CRR for the examined values of wind speed is shown in Sirajuddin and Gettu  examinedw/c ratios is shown in 3 of RTSF eliminated cracks at low and mid w/c ratios, while 40 kg/m3 of RTSF eliminated the cracks at the high w/c ratio.The CRR for various w/c ratio of 0.5 and observed that the fibres were effective in reducing the crack width, but the cracks were still visible at the surface of the concrete.Mazzoli et al.  examinedw/c ratios and dosages of RTSF were used to assess their effect in restraining cracking in fresh concrete. From the results, the following conclusions can be drawn:Harsher environmental conditions result in earlier and more severe cracking .Temperature is the most important parameter in accelerating cracking, whilst wind speed is the least significant.The use of higher water/cement ratios increases cracking, primarily as a result of earlier and heavier bleeding.3 eliminating the cracks completely in the harshest of the examined environments.The addition of RTSF fibers was found to be effective in controlling plastic shrinkage cracking, with 40 kg/mThis paper discusses the impact of low, mid, and high environmental conditions on the plastic shrinkage of concrete. Two


The recent interest in 3D printing with concrete has generated great interest on how inhomogeneities arise and affect performance parameters, in particular strength and durability. With respect to durability, of particular interest is how 3D‐printed layer interfaces can impact transport of species of interest, such as moisture, chlorides or carbon dioxide in carbonation processes. This is of particular interest considering that the primary use case of 3D‐printed concrete has been as a lost formwork for a cast structural concrete, and thus it is of interest to determine the carbonation resistance. This study consists of a preliminary look at the microstructure after accelerated carbonation of a 3D‐printed concrete used as a lost formwork. Preferential carbonation is observed in the layer interfaces compared to the bulk of the printed filaments, possibly related to porosity from air voids or a locally high capillary porosity corresponding to the lubrication layer. The concrete of this study is also printed through use of a set accelerator, with a low viscosity and low initial yield stress but a rapid buildability rate. Samples were produced as they would be expected to be used, that is through the use of 3D‐printed concrete as a formwork for a cast concrete. Good structural engineering design would demand the placement of reinforcement bars as close to the edge of the structural element as possible (where tension is highest), but design codes also demand a concrete cover to mitigate corrosion by carbonation. Therefore, it is worth examining if printed concrete could conceivably be used as a cover layer in 3D‐printed components, and our analysis of this 3D‐printed sample is made with this in mind.22.12.1.1w/b of 0.39 and sand to binder ratio of 1.9, before eventually being mixed in the print nozzle with an accelerator paste consisting of calcium aluminate cement (CAC) which is also retarded with sodium gluconate, at an approximate dosage of 10% OPC substitution.The mix used for 3D printing has also been reported elsewhere.2.1.2w/c = 0.6 and maximum grain size of 16 mm.The mix for the cast concrete was a standard C20/25 mix with a CEM I 42.5 N (Holcim Normo 4) at 2.2Printed samples were produced by the 3D printing process at ETH Zurich also described in.2 gas bottle. The CO2 level was maintained between 50% and 90% at a RH of 70% for a period of 2 months before the samples were removed.The samples were carbonated in accelerated conditions in a self‐made carbonation chamber made of a sealed box connected to a CO2.3The produced samples were then cut and immediately had phenolphthalein and thymolpthalein applied to check for carbonation depth. Regions of interest were identified and a polished thin section was produced from layers approximately middle height (as there was no major variation observed in carbonation behaviour from top to bottom) to observe the microstructure with both optical and scanning electron microscopes.2.4A stereo microscope Leica M 60 was used for the observation of bulk samples and a Leica DM750 P in both plane and polarised light mode was used for observing the thin section. SEM‐EDX observation (BSE images) and point analyses were performed with a QUANTA 2003D and EDX system from EDAX company.3Figure In Figure Figure Figure 4At the cast–print interface, one can observe a separation between the printed and cast layers, which could be due to drying shrinkage differences between the two materials and their different production times. The uniform carbonation depth of 0.5 mm in the printed concrete is likely due to the 10 days delay between the printing and the casting, however, because no corresponding carbonation depth was observed in the cast concrete. It is also worth noting that the highly advanced degree of hydration in this zone could be due to sorption of water from the cast concrete. Overall, examination of this region highlights the importance of understanding the relationship between printed formworks and cast concrete in terms of the overall process and performance.In the ‘transition zone’ (Zones 2 and 3), it is quite apparent that carbonation is occurring nonuniformly within the printed concrete, and that it is occurring with greater speed within the layer interfaces, indicative of a difference in the diffusivity between the layer interfaces and the bulk. This higher carbonation speed in the layer interfaces contributes to faster carbonation in the bulk of the print filament as well, essentially by allowing carbonation from the top and bottom of the filament and effectively reducing the diffusion length.It is interesting to note that the thickness of the carbonation layer at the layer interface is approximately 1 mm. It is well known that a ‘lubrication layer’ forms during pumping of concrete, in which a paste rich zone forms at the wall of the pipe,As noted earlier, most studies to date have focused on presence of defects and air voids in the interlayer, with a primary focus on their effect on the strength. Numerous X‐ray micro computerised tomography studies have detailed their presence and extent with respect to processing parameters such as printhead speed, pressure, and height, layer interval time, and ambient conditions with respect to moisture,w/c ratio. The same physical principle that leads to a depletion of aggregate next to the wall will also lead to a depletion of binder particles, which would lead to an effective increase in the w/c ratio within a certain range of the layer interface. This moisture content is enhanced by pressurised bleeding, noted by Sanjayan et al.w/c ratio would manifest itself in a higher degree of hydration, which was observed in this sample, across the width of the layer interface. This is corroborated by Van Der Putten et al.,w/c at the layer interface would lead to a higher capillary porosity and thus enhanced transport.Capillary porosity is strongly influenced by the w/c and concomitantly higher capillary porosity would allow portlandite to precipitate there preferentially, thus also to carbonate there preferentially.The dense carbonation layer at the layer interface could also be indirect evidence of this locally high capillary porosity. One would expect that carbonation products would form generally where there is a high amount of portlandite, such as in the traditional interfacial transition zone (ITZ). For example, Shen et al. demonstrated that carbonation products formed significantly in the ITZ.2 transport.A final discussion point could be made regarding the high degree of microcracking observed. While this could easily be related to sample preparation, it could also bear some relation to material, processing, and shrinkage. The printed material is essentially a binary OPC‐CAC system, relying on ettringite formation to develop rapid buildability strength.5The 3D‐printed concrete of this study showed nonuniform carbonation due to the presence of the layer interfaces, with the layer interfaces showing faster carbonation. This can be from enhanced transport due to either increased air content or capillary porosity in the layer interface. The implications for this require further study, as this would negatively impact the possibility of using a printed formwork as a cover layer.The interface of a structural concrete cast into a 3D‐printed formwork (‘cast–print’ interface) requires attention in future studies, especially as this represents the predominant use case of this technology.A more complete study is needed regarding the generation and evolution of microstructural heterogeneities in 3D‐printed concrete with respect to carbonation, in particular with respect to the capillary porosity at the layer interfaces.The primary conclusions and recommendations that can be drawn from this preliminary look at the microstructure of carbonated 3D‐printed concrete are the following: