Steel reinforced grout systems for the strengthening of masonry structures

Accepted Manuscript Steel reinforced grout systems for the strengthening of masonry structures Stefano De Santis, Gianmarco de Felice PII: DOI: Reference:

S0263-8223(15)00787-4 http://dx.doi.org/10.1016/j.compstruct.2015.08.094 COST 6799

To appear in:

Composite Structures

Please cite this article as: De Santis, S., de Felice, G., Steel reinforced grout systems for the strengthening of masonry structures, Composite Structures (2015), doi: http://dx.doi.org/10.1016/j.compstruct.2015.08.094

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COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

Steel reinforced grout systems for the strengthening of masonry structures Stefano De Santisa, Gianmarco de Feliceb,* Roma Tre University, Department of Engineering. Via Vito Volterra 62, 00146 Rome, Italy. a

[email protected], b [email protected] * Corresponding author. T: +39 06 5733 6268. F: +39 06 5733 3441.

ABSTRACT Steel Reinforced Grout (SRG) systems, comprising Ultra High Tensile Strength Steel cords embedded in mortar matrix, are increasingly used for the strengthening of existing structures. Their mechanical properties, however, still need to be fully investigated in order to provide reliable values of tensile and bond strength, strain and stiffness for design purposes. With this aim, four SRG systems made of two textiles combined with two mortars were tested. The durability of the textiles, the tensile behaviour of the composites and the shear bond performance on different substrates, including historic and modern bricks and tuff units, were investigated. The role of the layout of the textile, the characteristics of the matrix and the properties of the substrate are discussed. Finally, test results are combined to derive engineering design parameters for different field applications. Keywords Steel Reinforced Grout (SRG); Durability; Tensile behaviour; Shear bond performance; Strengthening; Qualification.

1.

INTRODUCTION

Composite materials have been increasingly used in the last two decades for repair and strengthening reinforced concrete and masonry structures. They consist of high strength textiles externally bonded to the surface of the structural elements and, thanks to their high strength-to-weight ratio, provide a significant improvement of the structural capacity with minimum mass increase. Nevertheless, drawbacks related to brittle failure, long-term durability, sensitivity to impacts, and high cost have limited their widespread use. After the diffusion of Fibre Reinforced Polymers (FRPs), mortar-based composites have been recently developed, which make use of high performance textiles (continuum dry fibres arranged in the form of open mesh or fabric) externally bonded with either cement or lime mortars. These systems are known as either TRMs (Textile Reinforced Mortars) or FRCMs (Fibre Reinforced Cementitious Matrices). Despite their adhesion strength may be lower than FRPs in certain cases, they offer important advantages in terms of fire resistance, vapour permeability, removability, and ease, time and cost of installation, thanks to the use of inorganic matrices in place of resins [1]. Numerous textile materials can be used, including carbon [2,3], glass [4,5], steel [6-8], basalt [9], PBO [1012], and natural fibres [13,14]. Among all these possibilities, steel-based reinforcements offer particularly good mechanical performance thanks to the high tensile strength of the textile and the effective cord-tomortar interlocking [15,16], at relatively low costs. Steel textiles are constituted by cords of Ultra High Tensile Strength Steel (UHTSS) micro wires and have been initially developed for the reinforcement of automobile tyres. The first application in civil engineering was proposed for the flexural strengthening of reinforced concrete beams [17] taking the name of Steel Reinforced Grout (SRG). Then, SRG has been used in other research studies on reinforced concrete beams [18], masonry arches [19] and walls [20,21]. The need of effective, versatile and cost efficient strengthening methods has encouraged producers to develop and sell steel textiles for structural rehabilitation purposes, such that SRG has already been used in the field for the seismic retrofitting of masonry walls [22] and vaults [23]. Based on information provided by the suppliers, the steel cords available today in the market have tensile strength of 2800-3200N/mm2 and Young’s modulus of 180-210kN/mm2 . Depending on the spacing between the cords, the surface mass density ranges widely from 600g/m2 to 3300g/m2 corresponding to maximum loads per unit width varying from 230kN/m to Page 1 of 19

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1300kN/m. Given their small diameter (0.1-0.5mm), the wires are either coated (with zinc or brass) or made of stainless steel, to provide protection against corrosion. Despite the scientific and industrial research carried out so date, and the existing applications realized, there are still important issues which need to be further investigated, such as the influence of textile layout, mortar strength and substrate properties on the SRG-to-masonry bond performance and the durability of the steel textiles. The former is crucial for the effectiveness of a broad range of strengthening works, while the latter ensures the long-term protection of the reinforced structure. Moreover, grounded experimental data sets are necessary to provide practitioners with mechanical properties of SRG systems for design purposes. This paper presents a laboratory investigation carried on SRG composites for the strengthening of masonry structures, which provides a mechanical characterization leading to the identification of design parameters. Aiming at reproducing the variability of available textiles and mortars, four SRG systems were tested, combining two textiles (differing for the density, namely 12 and 4 cords/inch) with two matrices (a mineral mortar and a natural hydraulic lime mortar), as described in Section 2. First, tensile tests were carried out on textile specimens to derive mechanical properties and durability against salt attack and mechanical flexure (Section 3). Second, direct tensile tests were carried out on composite specimens to characterize the entire response under tension, from the un-cracked stage up to tensile rupture, and derive information on crack pattern (Section 4). Third, shear bond tests were performed on two types of modern bricks (a strong one and a weak one), historic bricks and tuff units, to investigate the SRG-to-substrate load transfer performance. The influence of the properties of textiles, matrices and substrates on bond strength and failure mode are analysed (Section 5). Finally, test results are combined to identify engineering design parameters for different field applications (Section 6).

2. 2.1.

PROPERTIES OF THE MATERIALS AND EXPERIMENTAL PROGRAMME Materials

Unidirectional textiles of Ultra High Tensile Strength Steel (UHTSS) cords (Fig. 1a) were tested. Cords are made out of five wires with 0.11mm2 cross section area each, three rectilinear and two twisted around them at a short lay length to enhance the interlocking with the mortar. Wires are galvanized (coated with zinc) to provide protection against rusting, and are installed on a supporting glass mesh. Two different textiles were tested, having density of 12 and 4 cords/inch. In the former (S12), the design thickness is 0.254mm and the mass density is 2000g/m2 (Fig. 1b). Cords are placed adjacent to each other two by two such that the spacing between couples of cords is 4.25mm, in order to allow for the protrusion of the mortar during installation and to promote the matrix-to-textile interlocking. In the latter (S4), cords are equally spaced 6.35mm to obtain a textile with 0.084mm design thickness and 670g/m2 mass density (Fig. 1c). Composite specimens were manufactured with two mortars, such as a mineral mortar (identified by letter M) with a binder of natural kaolin and bauxite, and a mineral-NHL mortar (L), comprising natural kaolin, bauxite and hydraulic lime binders. Their compressive strength (fcm), Young’s modulus (Ecm), tensile strength (ftm), and grain size range (D) are collected in Table 1. Four SRG reinforcements were tested by combining the two textiles and the two matrices. Each system is identified by its constituents (textile and mortar). For instance, the composite comprising the textile with 12 cords/inch and the mineral mortar is named S12M. The bond performance of these four SRG systems was tested on four substrates, such as strong and weak modern bricks, historic bricks and tuff units, whose properties are collected in Table 1, also reporting the Coefficients of Variation (CV) in round brackets (more details on substrates will be provided in Section 5). 2.2.

Experimental plan

Aiming at characterizing the SRG reinforcements, the following three tests were carried out: (i) Tensile tests on textile specimens, to derive strength, strain corresponding to peak stress, and tensile modulus of elasticity. The durability against salt attack (to which the textile may be exposed due to mortar cracking) was also investigated after artificial ageing in substitute ocean water. Moreover, having in mind confinement applications, the deterioration induced by mechanical flexure (both by itself and in combination with salt attack) was investigated. (ii) Tensile tests on SRG specimens, which provided the whole stress-strain response and the main mechanical parameters of the three response stages of mortar-based composites under tension. These are (I) un-cracked stage (in which the mortar matrix contributes to both load bearing capacity and stiffness), (II) crack development stage (during which crack pattern develops progressively), and (III) Page 2 of 19

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cracked stage (in which crack pattern is completely developed). The parameters of stages I and II are the Young’s modules, the transition points between one stage and the following one, and the crack width and distribution, which are expected to affect the durability of the textile exposed to the external environment. The parameters of stage III are the Young’s modulus and the ultimate stress with the corresponding strain, and may be decisive for the structural applications in which the strength of the textile can be fully exploited (e.g., the confinement of columns and the extrados reinforcement of vaults with mechanical end anchors/pivots). (iii) Tests on the reinforcement-to-substrate shear bond performance, which provided the maximum load that can be transferred from the structure to the strengthening system, the corresponding slip (i.e., the relative displacement between reinforcement and substrate), and the failure mode. Recent studies [7,11,24] have shown that failure modes in mortar-based externally bonded reinforcements may significantly differ from those observed on FRPs, in which debonding generally involves the substrate. Due to the lower strength of the mortar matrix if compared to resin, failure in FRCMs may occur also on the substrate-to-reinforcement or on the matrix-to-textile interfaces, or also by textile slippage. The bond performance affects the effectiveness of the strengthening work for those structural applications in which the load is transferred by shear, such as the reinforcement of masonry walls towards both in-plane and out-of-plane loads, the strengthening of vaults and the strengthening of reinforced concrete beams in bending and/or in shear. All tests were carried out using a Material Testing Systems (MTS) load frame (Fig. 2). Load was applied by a 500kN hydraulic actuator under displacement control (machine compliance <0.05%), and recorded by a load cell with 0.2% accuracy and 0.01kN resolution. Due to the possible occurrence of brittle phenomena, different displacement rates were selected for the different test types, such as 0.02mm/s for tensile tests on textile specimens, 0.01mm/s for tensile tests on composite specimens, and, finally, 0.003mm/s for shear bond tests. Data were acquired at 10Hz sampling frequency by means of a National Instruments NI PCI 6281 Multifunction Data Acquisition (DAQ) system, provided with three units NI SCXI 1520 Universal StrainGauge Input Module, with eight Channels each; the acquisition software was developed in LabView environment.

3. 3.1.

MECHANICAL PROPERTIES AND DURABILITY OF STEEL TEXTILES Testing setup

The mechanical properties of the steel textiles were derived through tensile tests on specimens having 600mm length and 50.8mm width, made out of either 24 (for S12) or 8 (for S4) cords (Fig. 2a). Aluminium tabs (90mm×55mm×3mm) were glued on the ends of the specimens by means of a strong structural adhesive to ensure uniform stress distribution and absence of sliding in the gripping areas (Fig. 2b). A careful smoothing of the tabs and a sufficient amount of adhesive were necessary to provide an adequate contact with the textile and prevent its premature rupture due to local stress concentrations [16]. Global displacements were acquired by a Linear Variable Differential Transformer (LVDT) integrated in the testing machine, recording the relative displacement from end plate to end plate with 0.05% accuracy and 1µm resolution. Strains were derived as the recorded displacement divided by the initial distance between the clamping wedges. In order to confirm the reliability of global measures and detect possible sliding in the gripping areas, strains were also detected by an MTS extensometer having 50mm gage length, +25/-5mm range, 0.18% accuracy and 10µε resolution, placed in the middle of the specimens (Fig. 2a). Stresses were computed as the recorded load divided by the cross section area of the textile, which corresponds to the product of design thickness and specimen width. 3.2.

Artificial ageing and test results

First, tensile tests were carried out on six specimens of S12 and six of S4. The response was characterized by an initial linear elastic phase, followed by a nonlinear one, preceding the failure, which took place as the nearly simultaneous rupture of the cords (Fig. 3). The average tensile strength (ff) was 3082N/mm2 for S12 and 3186N/mm2 for S4 (Fig. 3a). Beyond the physiological scatter of experimental data, the slightly lower peak stress of S12 with respect to S4 may be due to the non-uniform load distribution among the cords, because of unavoidable (howsoever small) misalignments that may occur while clamping. The tensile modulus of elasticity (E f), computed between 30% and 60% of ff, was 183kN/mm2 for both textiles. The maximum attainable load per unit width (Ff) and the strain corresponding to the peak stress (εf) are collected in Table 2 (CV in round brackets) together with ff and Ef. Page 3 of 19

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The extremely small diameter of the steel wires makes the durability against rusting a crucial issue for the effectiveness of the strengthening system in the long-term. In order to investigate the durability of textiles against external aggression, specimens were aged in Substitute Ocean Water (SOW) solution, prepared according to ASTM D1141 US standard [25], and tested under tension to quantify the deterioration induced by salt attack. Different ageing durations were selected, such as 15, 30 and 41 days, the latter corresponding to 1000 hours. Then, specimens were washed with pure water, the aluminium tabs were applied and the tests were carried out. The stress-strain response curves were nearly superposed to the reference ones (Fig. 3b). After 41 days of ageing, a strength deterioration of 4.5% (Fig. 4a) and a reduction of the tensile modulus of elasticity of about 7% (Fig. 4b) were found. The coating layer locally deteriorated, as revealed by the Scanning Electron Microscope (SEM) analysis (Fig. 5a), but proved to effectively protect the steel from heavy corrosion. The experimental outcomes comply with the US standard for the qualification of FRCM reinforcements [26], which requires a maximum reduction of strength of 15% after 1000h immersion in Substitute Ocean Water. It should also be considered that the durability tests were carried out on dry textile specimens, while in structural applications the steel cords are embedded in the mortar matrix, which, although cracked, could still provide a certain protection against salt attack. Nevertheless, the possibility of extrapolating these results to longer exposure durations, despite beyond the aims of this paper, still needs to be investigated and, to this purpose, tests after 3000h ageing could be carried out, as recommended by [26]. Thanks to the high tensile strength and stiffness of steel textiles, confinement of columns and pillars is a promising field of application of SRG. In this case, the installation requires that the steel cords are flexed at the corners to ensure adhesion. Aiming at investigating the possible deterioration induced by flexure, tensile tests were carried out on S4 specimens flexed at 90°. A reduction of about 15% and 13% was found for tensile strength and modulus of elasticity, respectively (Fig. 3c and Fig. 4). It should be noted, however, that for confinement applications design guidelines generally limit the ultimate strain of the textile, e.g., at 0.4% [27] (that is well below the strain at failure resulting from experiments), in order to preserve the shear strength of the confined element. Finally, tensile tests were carried out on S4 specimens that had been first flexed at 90° and then aged for 41 days in Substitute Ocean Water, to study the combined effect of both flexure and salt attack. The stress-strain response curves are represented in Fig. 3d, and an overall deterioration of nearly 20% and 15% was found for ff (Fig. 4a) and Ef (Fig. 4b), respectively, which approximately corresponds to the sum of the reductions observed after flexure and ageing separately. Flexure produced a non-negligible mechanical deterioration, but did not strongly damage the protection layer (Fig. 5b), which was then locally corroded by salt attack (Fig. 5c) but continued protecting the steel wires from heavy rusting.

4. 4.1.

TENSILE BEHAVIOUR OF SRG STRENGHTENING SYSTEMS Testing setup

In order to characterize the tensile response of SRG systems, prismatic composite specimens were manufactured with the two textiles (S12 and S4) and the two mortars (M and L). Specimens had 500mm total length and 50mm×10mm cross section, and comprised 24 cords for S12 and 8 cords for S4. Prisms were casted in Plexiglas moulds, cured at 95% R.H. for 24 hours before demoulding, and for the following 27 days in water, as recommended for the qualification of repair mortars [28]. Finally, they were left for 7 days in laboratory conditions (18-20°C temperature and 50-60% R.H.) before testing (Fig. 2c). By doing so, shrinkage induced deflection and cracking of the specimens are avoided. Proper manufacturing, wellcontrolled curing conditions and careful storage and handling of the specimens may heavily affect the dispersion of test results, and the uncertainties related to all these factors have not been fully investigated yet. Clearly, the curing conditions undergone by SRG reinforcements in real field applications differ from those, standardized and well-controlled, followed in the laboratory. The impossibility of creating a water-saturated curing environment in the construction site is however partially balanced by the larger size of the casting and by the presence of the substrate (if sufficiently wetted before installation), which limits shrinkage development. Two layers of GFRP reinforcement (glass fibre textile bonded with epoxy resin) were applied around the ends of the specimens for a length of 110mm to ensure a uniform load distribution and avoid mortar crushing in the vicinity of the gripping areas [16]; the gripping length was 90mm (Fig. 2d). Five monotonic tests were carried out on each SRG system. In order to record displacements and strains, in addition to the LVDT of the testing machine and the extensometer, two linear potentiometers (±12.5mm range, 0.15% accuracy, and 10Page 4 of 19

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mm resolution) were also applied on the mortar by means of aluminium plates, with a base length of about 250mm (Fig. 2c). Test results are expressed as stress-strain relationships. Strains were evaluated either from the extensometer (before cracking), or from the LVDT and the potentiometers, dividing the recorded displacements by the measurement base of the device. The comparison between global strains from LVDT and potentiometers allowed ruling out sliding occurrence in the gripping areas. Stresses were derived as the applied load divided by the cross section area of the textile, to prevent the results from being affected by the variations of mortar cross section, which are unavoidable especially in field applications. 4.2.

Test results

The response curves of SRG specimens under tension are shown in Fig. 6, the mean values related to Stages I and II (un-cracked and crack development), are collected in Table 3, while those related to Stage III (cracked stage up to failure) are reported in Table 4. In the tables, σI and εI are the stress and the strain of the transition points between stages I and II, while σII, and εII identify the transition between Stages II and III; ft and Ft are the peak stress and the corresponding load per unit width, εt is the peak strain, Et,I, Et,II and Et,III are the tensile modules of elasticity of the three stages; finally, δ is the saturation crack spacing, i.e., the mean distance between cracks in the last stage (in which crack pattern has stabilized). In all cases, the response was mainly governed by stage III, indicating that mortar cracked relatively soon and crack widening was the prevalent mechanism. Failure occurred by tensile rupture of the cords, after a loss of linearity (due to a certain ductility of the steel). Both the peak stress and the tensile modulus of elasticity in the last stage were comparable to those of dry textiles (Table 4). SRG systems with S12, however, showed a slight strength reduction, due to stress concentrations near the cracks (where the transverse redistribution effect of the mortar vanishes), leading some lateral cords to reach rupture before the others. This is confirmed by the stress drops in the response curves, well before overall failure. The peak strain of the composites was similar to that of the textiles, except for S4M (4 cords/inch textile with mineral mortar), in which the stiffening effect of the matrix caused a reduction from εf=2.26% (textile) to εt =2.06% (composite). The first response stages were clearly visible in the stress-strain curves of SRG systems with the stronger matrix (S12M and S4M series, Figs. 6a and 6c, respectively), as the un-cracked mortar significantly contributed to both stiffness and load bearing capacity. Both the tensile modulus of elasticity (Et,I) and the stress value at the first transition point (σI) were comparable (note that stresses are referred to the crosssection area of the textile, which in S12 is three times larger than in S4). The use of the weaker NHL mortar in place of the mineral mortar led to remarkable reduction of both the tensile modules of elasticity in the first two stages and the stress values of the transition points (Table 3), such that stages I and II are hardly identifiable in the response curves (Figs. 6b and 6d). Sliding of the textile within the mortar was never observed, differently from other mortar-based composites [24,29]. All SRG system displayed transverse cracks regularly spaced (Fig. 7). The average saturation crack spacing (δ) varied from 19.5mm (S4L, Fig. 7d) to 24.4mm (S12M, Fig. 7a), corresponding to average crack width (at failure) in the 0.38-0.48mm range. S12 specimens displayed a slightly longer distance between cracks than S4, indicating a relatively lower textile-to-mortar stress transfer capacity, due to the fact that the protrusion of the mortar is more difficult through the voids of a denser textile. The use of a stiffer matrix (M instead of L) also led to a slightly higher crack spacing. Beyond these small differences, both textiles demonstrated an extremely good interlocking, not only within the stronger mineral mortar but also within the weaker NHL mortar, thanks to the shape of the steel cords, and particularly to the two external wires twisted around the three rectilinear ones. In principle, the relationships developed for reinforced concrete members under tension/bending [30] can be extended to represent cracking in SRG composites. According to this approach, the larger is the cross section area of steel reinforcement and the better is the concrete-to-rebar bond (plain/high bond rebars), the closer and narrower are, on average, the cracks. Based on the test results achieved in this study, in SRG the stress transfer performance between cords and mortar plays a much more important role than the cord density (closer cracks appeared on composites with S4, providing a better interlocking with the mortar thanks to the larger cord spacing). Moreover, if the paired cords in S12 are imagined as one single larger cord, the increase of crack spacing and width is consistent with the dependence on steel rebar diameter assumed in reinforced concrete.

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5. 5.1.

SHEAR BOND PERFORMANCE Substrates and preparation of the specimens

Bond tests were carried out on the following substrates: • Modern strong bricks (identified by acronym SB), measuring 120mm×240mm×55mm; • Modern weak bricks (WB), measuring 120mm×250mm×55mm; • Historic bricks (HB), produced in 1920s, measuring 120mm×255mm×55mm; • Tuff units (TU), measuring 175mm×245mm×110mm. The mechanical properties of the substrates are listed in Table 1. The surface of the modern bricks was roughened with a bush hammer to promote the adhesion of the reinforcement. On the contrary, a minimum surface preparation was made on the historic bricks, consisting of brushing and cleaning with compressed air, to accomplish the preservation requirements faced when dealing with cultural heritage. For the same reason, only SRG systems with lime mortar were tested on this substrate, while both mineral and lime mortars were applied to the other substrates. Tuff units were treated with an aqueous solution of potassium silicate, applied prior to the laying of the first mortar layer, in order to consolidate the tuff and remove dust. A standard wet lay-up procedure was followed to bond the textiles on the substrates. After substrate preparation, a first 5mm thick layer of mortar was applied with the help of aluminium frameworks. The textile was then placed by hand and pressed slightly into the fresh mortar, which protruded through the voids between steel cords. The textile strip comprised 24 cords for S12 and 8 cords for S4, and was 600mm long. The bonded area was 200mm long and 50mm wide. Finally, a 5mm thick layer of mortar was applied on top. Specimens were kept wet (R.H.>95%) for the first 3 days, and then stocked for other 25 days in standard laboratory conditions. In shear bond tests, even more than in direct tensile tests, the manufacturing and curing of specimens may largely affect the results. More specifically, an insufficient amount of water in the substrate or in the matrix during installation and curing may compromise the setting process of the mortar and, consequently, a decrease in its mechanical properties (especially its adhesion to the substrate). Table 5 provides an overview of the shear bond tests. Specimens are labelled using the acronym of the SRG system followed by that of the substrate. For instance, the series of 4 cords/inch textile applied with lime mortar to modern strong bricks is named S4L-SB. Five tests were carried out on each series, for a total of 70 tests. 5.2.

Testing setup

A single-lap setup was used. The specimen was placed on a supporting steel device, made out of 10mm thick steel plates welded to form an angle of 90°, and appropriately stiffened to avoid undesirable distortions (Fig. 8a). The steel frame was held from above by the wedges of the testing machine, while the reinforcement strip was clamped from below (Fig. 8b). Smoothed aluminium tabs were glued to the end of the unbonded textile to ensure adequate clamping. In order to guarantee load balance, the specimen was also blocked from behind by steel C-plates connected through ∅20 threaded bars to the frame. The device was designed to align the upper plate of the frame and the textile strip, thus ensuring that a pure shear load was applied on the reinforcement and avoiding parasitic normal stresses on the substrate-to-matrix interface. Two linear potentiometers (±5mm range, 0.05% accuracy, <10-4mm resolution) were installed to record the relative displacement (slip) between substrate and textile, in its first unbonded section on the loaded end. Additionally, an MTS extensometer (50mm gage length, +25/-5mm range, 0.18% accuracy and <10-5 resolution) was placed on the unbonded textile to record strains (Fig. 8a). 5.3.

Test results: shear bond strength and failure modes

The response curves of shear bond test are shown in Fig. 9 for SRG systems with S12 and in Fig. 10 for those with S4. The plots have the slip on the x-axis and the force per unit width (of the textile) on the y-axis. Each graph shows two series, which differ only for the mortar. The average values and the coefficients of variation of maximum force (Fd) and corresponding slip (sd) are collected in Table 6, together with the stress in the textile at debonding (σd) and the exploitation ratio of textile strength (σd/ff). As a general trend, the response of all the tested SRG systems was characterized by an initial stiff behaviour, followed by a reduction of stiffness, and a brittle failure. Despite a relatively wide variation of results among test series, a higher stiffness was related to a stiffer textile (S12 instead of S4) and a stiffer matrix (mineral instead of lime mortar). The slip corresponding to the peak load ranged between 0.10mm to 0.34mm for SRG systems with S12 (Fig. 9) and between 0.25mm and 1.76mm with S4 (Fig. 10). Given the relatively low Young’s modulus of the substrate, the slip recorded on tuff units was larger than that on the bricks for both textiles (Figs. 9d and 10d). Page 6 of 19

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Both the failure mode and the maximum load depended on the density of the textile, the strength of the mortar, and the surface roughness of the substrate. Three failure modes were observed such as cohesive debonding involving the substrate (failure mode A), debonding at the mortar-to-substrate interface (B), and debonding at the textile-to-mortar interface (C), as illustrated in Fig. 11. The good cord-to-mortar interlocking impeded the sliding of the textile within the matrix, and the high tensile strength of the steel cords avoided the tensile rupture. Finally, since the textiles were not made by bundles of wires, the telescopic rupture that has been observed with uncoated carbon, glass, PBO and basalt fabrics [10,11,24] could not take place. Table 6 reports the prevalent failure mode for each series, while the failure modes occurred in all tests are reported in Fig. 12 by varying the shape of the marker. Failure mode A was observed in SRG systems with the low density textile bonded with the mineral mortar (S4M) to weak bricks (S4M-WB series, Fig. 11a) and tuff units (S4M-TU, Fig. 13), due to the relatively weakness of the substrates. The strength of the mortar and the roughness of the substrate (either artificially improved on the brick, or due to the natural porosity of the tuff) prevented the detachment at the substrate-tomatrix interface (mode B). Similarly, the good textile-to-matrix interlocking impeded the detachment between textile and matrix (mode C). On brick substrate, the mean bond strength was Fd=115kN/m corresponding to a stress in the textile of σd=1365N/mm2 (43% of ff), while on tuff Fd=158kN/m, σd=1876N/mm2 = 59% ff (Table 6). The same SRG system applied to the stronger brick (S4M-SB series) failed by detachment at the mortar-tobrick interface (mode B). Debonding within the substrate was impeded by its strength and a maximum load higher than on weak bricks was reached (Fd=142kN/m, σd=1688N/mm2 = 53% ff). The SRG systems with S4 textile and lime mortar mainly failed by detachment at the textile-to-matrix interface (mode C), indicating that the weaker is the matrix the lower is the load transfer performance with the textile (Fd was 106kN/m on strong brick, 69kN/m on weak brick, and 141kN/m on tuff). The only exception is represented by S4L-HB series, in which failure occurred at the mortar-to-substrate interface, as a result of the minimum surface preparation. In this case, the bond strength was 112kN/m (σd=1329N/mm2 = 42% ff). The shear bond strength of S12M systems was between 98kN/m (on strong brick) and 123kN/m (on tuff), while that of S12L systems was between 68.4kN/m and 77.2kN/m on bricks and 113.8kN/m on tuff. The smaller cord spacing in S12 textile with respect to S4 was responsible for a lower textile-to-mortar load transfer performance, as revealed by the prevalent occurrence of failure mode C. As a consequence, the bond strength was lower than that of the corresponding systems with S4 (i.e., with the same matrix and on the same substrate) on modern strong bricks, historic bricks, and tuff units, while no significant difference was observed for modern weak bricks (Fig. 12). The stress in the textile at shear bond failure was between 12% and 15% of its tensile strength for S12M systems, and about 9-14% for S12L. SRG systems with S12 revealed lower exploitation ratios (σd/ff) than those with S4, as a consequence of the higher cord density, corresponding to a larger textile cross section and, given the same force, to a lower stress. The higher tensile strength of S12 could be exploited in applications to curved surfaces (e.g., extrados of arches or vaults), in which the compressive stresses (orthogonal to the textile) lead to higher failure loads than on straight substrates, as well as after installation of mechanical end anchors or pivots that may significantly increase the bond performances. The comparison between specimens with the same SRG system (e.g., S4L) indicates that the load transfer performance between reinforcement and substrate is influenced not only by the properties of the mortar and by the textile-to-matrix load transfer capacity, but also by the mechanical properties of the substrate and the roughness of its surface, even if this latter may not appear directly involved in the failure mechanism. Based on the results presented herein, the bond performance was 69kN/m on weak brick (S4L-WB), 106kN/m on strong brick (S4L-SB), and 141kN/m on tuff (S4L-TU). Similar trends were found for the other series, including those with S12 (see, for instance, the specimens with S12 and lime mortar). A possible explanation can be found by analysing the stress field in the mortar layer between textile and substrate. Fig. 13 represents a specimen from S4M-TU series. In this case, debonding occurred within the tuff substrate, allowing for a deep observation of the failure surface (Fig. 13a). The tensile load in the steel cords forms inclined struts in the underlying mortar layer (Figs. 13b and 13d), as revealed by the ribs in the failure surface (Fig. 13b) and by the transversal cracks on the side of the reinforcement (Fig. 13c). The component of the stress in the struts parallel to the SRG-to-substrate interface balances the external applied load. The maximum value that such parallel stress can reach depends on the combination of adhesion and interlocking, this latter relying on the roughness of the application surface. The higher is the roughness, the more inclined are the struts, leading to a higher parallel stress that can be reached before the orthogonal stress equals the tensile strength activating

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the bond failure. A wider experimental investigation and a refined micro-modelling study, which are beyond the scopes of this paper, would provide a deeper insight on this phenomenon.

6.

ENGINEERING DESIGN PARAMETERS

In this section, the results of the tests are combined to provide mechanical parameters for engineering design purposes. The characteristic strength values are determined as the mean minus 2.33 times the standard deviation, and then identified on the mean stress-strain curves of tensile tests to determine the corresponding strains (Fig. 14) [24]. The characteristic fractile factor of 2.33 is recommended by the design by testing procedure of Eurocode 0 for 5 tested specimens [31]. For structural applications of externally bonded SRG reinforcements in which tensile failure is expected to occur, the maximum attainable stress (σtk) is derived by reducing the characteristic strength of the unaged SRG composite under tension (ftk) to account for corrosion induced by salt attack. Such reduction can be expressed by means of a scalar coefficient ξ = ffaged / ffref, ffaged being the mean tensile strength after ageing in substitute ocean water, and ffref the mean tensile strength of unaged specimens. By doing so, the characteristic tensile strength is derived as σtk = ftk × ξ. The engineering design parameters (maximum stress σtk, corresponding strain εtk and maximum load per unit width Ftk) derived from the results of the present study are collected in Table 7, with ξ=0.955 (after 1000h ageing, Table 1). Thanks to the effective protection of zinc coating against rusting, and the low scatter of experimental data, the obtained strength values are relatively high, and range between 2506N/mm2 (corresponding to 636kN/m) for S12L and 2970N/mm2 (250kN/m) for S4L. However, despite the extremely reduced scatter, due to the small number of specimens tested, these numbers should be considered as preliminary results of a methodological proposal, rather than a quantitative conclusion. As for the stiffness, the average tensile modulus of elasticity in the cracked stage (Et,III) can be used, as listed in Table 4. In the applications in which the load is transferred from the structure to the reinforcement by shear/adhesion at the substrate-to-matrix interface, shear bond failure is expected. In this case, the design parameters can be identified through the qualification procedure proposed in [24], in which the results of shear bond tests and direct tensile tests on FRCM composites are combined. According to this method, the characteristic maximum stress (σdk) is determined on the base of shear bond tests, while the corresponding strain (εdk) is identified on the mean stress-strain response curve under tension (stresses are always referred to the cross section of the textile). Finally, the secant elastic modulus is calculated as Ed=σdk/εdk. In the present study, the characteristic strength values were generally close to the mean ones, given the limited dispersion of experimental data. Beyond the variability amongst systems and substrates, already discussed in the previous section, it is worth noting that the qualification points generally fell between stages II and III for SRG systems with S12 textile, and in stage III for those with S4 (Fig. 14), indicating that the effectiveness of the substrate-to-reinforcement stress transfer mechanism is not precluded by mortar cracking. The values of the secant stiffness corresponded or were slightly higher than those of the dry textiles. Since in SRG systems with S12 the qualification point fell in Stage II or at the beginning of stage III, the secant stiffness was higher than in those with S4 (240-400kN/mm2 for S12M and 193-215kN/mm2 for S12L versus 226-257kN/mm2 for S4M and about 180kN/mm2 for S4L, Table 7).

7.

CONCLUSIONS

The mechanical characterization of Steel Reinforced Grout (SRG) composites was carried out through tensile and shear bond tests. Four SRG systems were tested, combining two textiles (with 12 or 4 cords/inch) with two matrices (a mineral mortar and a natural hydraulic lime mortar). Tensile tests yielded to the following conclusions: • The unidirectional textiles with UHTSS cords displayed a tensile strength of about 3100N/mm2 and a Young’s modulus of 183kN/mm2. The corresponding maximum load per unit width is about 790kN/m for the textile with 12 cords/inch (S12) and 270kN/m for that with 4 cords/inch (S4). • Zinc coating effectively prevents rusting induced by salt attack. Specimens tested after 1000h ageing in Substitute Ocean Water displayed a reduction of about 5% for both tensile strength and Young’s modulus. Flexing the textiles produced a strength deterioration of 15% and a decrease in stiffness of 10%, but did not compromise protection against salt attack. Despite the low scatter of experimental data, a deeper investigation is still needed on more specimens and after longer ageing durations, since Page 8 of 19

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the durability of the steel cords (given their small diameter) is a fundamental requirement for the long-term effectiveness of the strengthening work. • The tensile behaviour of SRG systems displayed three characteristic response stages (as other mortarbased composites), such as (I) un-cracked, (II) crack development and (III) cracked. Test results indicated that the response of the composite mainly relies on the properties of the textile, as stage III is prevalent with respect to stages I and II. Indeed, the peak stress, the corresponding strain and the tensile modulus of elasticity in the cracked stage are very similar to those of the dry textile. However, the combination of a textile with lower density together with a stiff mortar may lead to an increase in stiffness due to the stiffening effect of the matrix; conversely, composites with a higher density may display a small decrease of tensile strength due to stress concentrations in the cracked sections. • Transverse cracks under tension developed at regular distance. The saturation crack spacing was about 20-25mm and the mean crack width at failure was 0.38-0.48mm. Slightly closer and narrower cracks developed in SRG systems with the more deformable mortar and the less dense textile. Beyond these small differences, both S12 and S4 textiles offered an extremely good interlocking within the mortar matrix (even with the relatively weak NHL mortar), thanks to the roughness of the cords, enhanced by the arrangement of five steel wires, and the fabric layouts, allowing for the protrusion of the mortar in the spacing between cords. Shear bond tests were performed on strong and weak modern bricks, historic bricks, and tuff units, providing the following results: • Three failure modes were observed. Cohesive debonding involving the substrate (A) occurred with the low-density textile applied with the stronger mortar to the weak bricks and tuff units, promoted by the weakness of the substrate material. Debonding at the textile-to-matrix interface (failure mode C) mainly occurred, especially with the weaker mortar and the denser textile, due to the relatively lower stress transfer capacity at the textile-to-matrix interface. Finally, applications to a stronger substrate, with smooth surface and no preparation, led to the detachment at the mortar-to-brick interface (B). • The shear bond strength varied widely depending on the density of the textile, the strength of the mortar, and the surface roughness of the substrate. Failure loads ranged between 68kN/m and 124kN/m for SRG systems with S12 textile and between 70kN/m and 157kN/m with S4. • The higher density of S12 textiles reduces the transfer capacity with the matrix and causes a lower bond strength than S4 in terms of load per unit width. Differently from FRPs, in which the shear bond performance depends on the cohesion between matrix and substrate, in SRG systems it also relies on the continuity of the mortar matrix in the cross section of the reinforcement, which, in its turn, depends on the protrusion of the mortar through the voids of the textile. • Beyond such difference in terms of maximum load per unit width, since the cross section area in S12 is three times larger than in S4, the stress values at failure are even lower. The exploitation ratio of the tensile strength of the textile varied between 9% and 16% for S12, and between 26% and 59% for S4. • The maximum loads reached on tuff substrates were higher than those on bricks, even when debonding occurred at the textile-to-matrix interface (mode C), suggesting that the load transfer performance between steel cords and mortar is influenced by the roughness of the substrate, even when this latter is not directly involved in the failure mechanism. The following design parameters were derived: • The combination of tensile tests on textile specimens after ageing and of composite specimens provided characteristic strength values comprised between 2506N/mm2 (corresponding to a load per unit width of Ftk=636kN/m and a strain of εtk=1.51%) for SRG systems with S12 and 2970N/mm2 (Ftk=250kN/m, εtk=1.77%) for those with S4. • Shear bond and direct tensile tests provided the characteristic shear strength (σdk), the corresponding strain in the textile (εdk) and the tensile modulus of elasticity (Ed) for each SRG system on each specific substrate. As for the systems with S12 textile, the characteristic shear strength varied between 172N/mm2 and 430N/mm2, corresponding to a load per unit width (Fdk) between 43kN/m and 109kN/m, εdk between 0.07% and 0.18% and Ed between 192kN/mm2 and 402kN/mm2. SRG reinforcements with the 4 cords/inch textile and the mineral mortar (S4M) displayed σdk=1284÷1720N/mm2, Fdk=107÷144kN/m, εdk=0.50÷0.76% and Ed =220÷247kN/mm2. Finally, when the lime mortar was used with the sparser textile (S4L) the design parameters resulted σdk=630÷1233N/mm2, Fdk=53÷104kN/m, εdk=0.35÷0.68% and Ed=180kN/mm2. Page 9 of 19

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•

The qualification point, identified by the coordinates (εdk, σdk) fell in stages II and III SRG with S12 textile, and in stage III with S4, indicating that the effectiveness of the substrate-to-reinforcement stress transfer mechanism is not precluded by the cracking of the matrix. Clearly, a larger number of tests should be carried out to perform more robust statistics and a probabilistic approach should be used to derive more reliable characteristic values. The experimental investigation presented in this paper showed the great potentialities of Steel Reinforced Grout systems for the repair and retrofitting of existing masonry constructions. Test results indicated that different solutions exist, which vary for the density of the textile and the properties of the mortar matrix, displaying different mechanical performances, in terms of maximum exploitable tensile load and reinforcement-to-substrate shear bond capacity. The tensile strength is proportional to the density of the steel cords, while the shear strength is promoted by the protrusion of the matrix in the spaces between adjacent cords. Therefore, if the tensile strength governs the design (e.g. for the confinement of columns or the extrados reinforcement of vaults with mechanical end anchors/pivots) denser textiles (e.g., S12) could be preferred, while if the bond capacity is the main design parameter a sparser textile (e.g., S4) seems to provide better performances. On the one hand, this indicates that best SRG system to be preferred in the strengthening design depends on the specific application, expected failure mode, strength of substrate material and characteristic of its surface, and need for fulfilling preservation criteria for historic masonry substrates. On the other hand, a further improvement of the structural performance of SRG reinforcements could be achieved by developing optimized systems, with appropriate density of the textile and suitable grain size range of the mortar matrix. The results of shear bond tests showed that the substrate surface needs to be prepared before the reinforcement is installed. Appropriate removal of dust and increase of roughness are fundamental to ensure a good shear bond performance, and prevent the occurrence of failure mode B. Additionally, and when feasible, the substrate should be consolidated to increase the reinforcement-to-substrate shear bond capacity. Finally, the surface of the substrate and the reinforcement need to be kept wet during mortar curing to ensure the effectiveness of the strengthening work. A sufficient amount of water in both the substrate and in the matrix during manufacturing, well-controlled curing conditions at high relative humidity, and careful storage and handling of the specimens are necessary to ensure the full development of the setting process of the mortar and avoid premature failure. The actual incidence of these factors on the reliability and dispersion of experimental results, especially in shear bond tests, would deserve a deeper investigation, in the perspective of defining standardized application and testing methodologies. The occurrence of different failure modes, involving either directly the substrate or the reinforcement system within its thickness, suggests that the theoretical approach (and the derived design relationships) developed for FRPs (which assumes cohesive debonding in the substrate) [27] cannot be directly extended to mortarbased composites. The mechanical properties of the mortar matrix, which play a fundamental role when failure occurs within the thickness of the reinforcement (mode C), should be included in the theoretical estimate of the maximum load.

ACKNOWLEDGEMENTS This work has been carried out under the research project ReLUIS-DPC 2014-2016, Thematic Area Innovative materials for interventions in seismic areas. Kerakoll S.p.A. provided materials and funding. The technical cooperation of Dr. Paolo Casadei is kindly acknowledged.

REFERENCES [1] Papanicolaou CG, Triantafillou TC, Karlos K, Papathanasiou M. Textile reinforced mortar (TRM) versus FRP as strengthening material of URM walls: in-plane cyclic loading. Mater Struct 2007;40(10):1081-1097. [2] D'Ambrisi A, Feo L, Focacci F. Experimental and analytical investigation on bond between Carbon-FRCM materials and masonry. Compos Part B-Eng 2013;46:15-20. DOI: 10.1016/j.compositesb.2012.10.018. [3] Carozzi FG, Colombi P, Poggi C. Calibration of end-debonding strength model for FRP-reinforced masonry. Compos Struct 2015;120:366-377. DOI: 10.1016/j.compstruct.2014.09.033. [4] Carozzi FG, Milani G, Poggi C. Mechanical properties and numerical modeling of Fabric Reinforced Cementitious Matrix (FRCM) systems for strengthening of masonry structures. Compos Struct 2014;107:711-25. DOI: 10.1016/j.compstruct.2013.08.026. [5] Gams M, Tomaževič M, Kwiecień A. Strengthening brick masonry by repointing - An experimental study. Key Eng Mater 2015;624:444-52. DOI: 10.4028/www.scientific.net/KEM.624.444.

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[6] Capozucca R. Experimental FRP/SRP-historic masonry delamination. Compos Struct 2010;92(4):891-903. DOI: 10.1016/j.compstruct.2009.09.029. [7] de Felice G, De Santis S, Garmendia L, Ghiassi B, Larrinaga P, Lourenço PB, Oliveira DV, Paolacci F, Papanicolaou CG. Mortar-based systems for externally bonded strengthening of masonry. Mater Struct 2014;47(12):2021-37. DOI: 10.1617/s11527-014-0360-1. [8] Malena M, de Felice G. Debonding of composites on a curved masonry substrate: Experimental results and analytical formulation. Compos Struct 2014;112(1):194-206. DOI: 10.1016/j.compstruct.2014.02.004. [9] Balsamo A, Di Ludovico M, Prota A, Manfredi G. Masonry walls strengthened with innovative composites. American Concr Inst, ACI Spec Publ 2011;2(275):769-86. DOI: 10.14359/51682454. [10] D'Ambrisi A, Focacci F, Caporale, A. Strengthening of masonry-unreinforced concrete railway bridges with PBO-FRCM materials. Compos Struct 2013;102:193-204. DOI: 10.1016/j.compstruct.2013.03.002. [11] D’Antino T, Carloni C, Sneed LH, Pellegrino C. Matrix-fiber bond behavior in PBO FRCM composites: A fracture mechanics approach. Eng Fract Mech 2014;117:94-111. DOI: 10.1016/j.engfracmech.2014.01.011. [12] Carozzi FG, Poggi C. Mechanical properties and debonding strength of Fabric Reinforced Cementitious Matrix (FRCM) systems for masonry strengthening. Compos Part B-Eng 2015;70:215-230. DOI: 10.1016/j.compositesb.2014.10.056. [13] Pacheco-Torgal F, Jalali S. Cementitious building materials reinforced with vegetable fibres: a review. Constr Build Mater 2011;25(2):575-81. DOI: 10.1016/j.conbuildmat.2010.07.024. [14] Codispoti R, Oliveira DV, Olivito RS, Lourenço PB, Fangueiro R. Mechanical performance of natural fiber-reinforced composites for the strengthening of masonry. Compos Part B-Eng 2015;77:74-83. DOI: 10.1016/j.compositesb.2015.03.021. [15] Razavizadeh A, Ghiassi B, Oliveira DV. Bond behavior of SRG-strengthened masonry units: Testing and numerical modeling. Constr Build Mater 2014;64:387-97. DOI: 10.1016/j.conbuildmat.2014.04.070. [16] De Santis S, de Felice G. Tensile behaviour of mortar-based composites for externally bonded reinforcement systems. Compos Part B-Eng 2015;68:401-13. DOI: 10.1016/j.compositesb.2014.09.011. [17] Huang X, Birman V, Nanni A, Tunis G. Properties and potential for application of steel reinforced polymer and steel reinforced grout composites. Compos Part B-Eng 2005;36:73-82. DOI: 10.1016/S1359-8368(03)00080-5. [18] Da Porto F, Stievanin E, Gabin E, Valluzzi MR. SRG application for structural strengthening of RC beams. ACI Special Publication 2012;286:119-32. DOI: 10.14359/51683908. [19] Borri A, Casadei P, Castori G, Hammond J. Strengthening of brick masonry arches with externally bonded steel reinforced composites. J Compos Constr 2009;13(6):468-475. DOI: 10.1061/(ASCE)CC.1943-5614.0000030. [20] Borri A, Castori G, Corradi M. Shear behavior of masonry panels strengthened by high strength steel cords. Constr Build Mater 2011;25(2):494-503. DOI: 10.1016/j.conbuildmat.2014.01.056. [21] De Santis S, De Canio G, de Felice G, Malena M, Mongelli M. Roselli I. Seismic performance of masonry walls retrofitted with steel reinforced grout. Submitted for publication. [22] Borri A, Castori G, Corradi M, Speranzini E. Shear behavior of unreinforced and reinforced masonry panels subjected to in situ diagonal compression tests. Constr Build Mater 2011;25(12):4403-4414. DOI: 10.1016/j.conbuildmat.2011.01.009. [23] Valluzzi MR, Modena C, de Felice G. Current practice and open issues in strengthening historical buildings with composites. Mater Struct 2014;47(12):1971-85. DOI: 10.1617/s11527-014-0359-7. [24] Ascione L, de Felice G, De Santis S. A qualification method for externally bonded Fibre Reinforced Cementitious Matrix (FRCM) reinforcement systems. Compos Part B-Eng 2015;78:497-506. DOI: 10.1016/j.compositesb.2015.03.079. [25] ASTM D1141-98(2013). Standard Practice for the Preparation of Substitute Ocean Water. ASTM International, West Conshohocken, PA, 2013. [26] ICC, International Code Council. AC434. Acceptance criteria for masonry and concrete strengthening using fiberreinforced cementitious matrix (FRCM) composite systems. ICC-Evaluation Service, Whittier, CA, 2013. [27] CNR, Italian Research Council. CNR-DT 200 R1/2013. Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Existing Structures. Italian Research Council, Italy, 2013. [28] CEN, European Committee for Standardization. EN 12190:1998. Products and systems for the protection and repair of concrete structures. Test methods. Determination of compressive strength of repair mortar. [29] De Santis S, de Felice G. Tensile behaviour and durability of mortar-based strengthening systems with glass-aramid textile. J Key Eng Mater 2015;624:346-53. DOI: 10.4028/www.scientific.net/KEM.624.346. [30] CEN, European Committee for Standardization. EN 1992-1-1:2004. Eurocode 2: Design of concrete structures - Part 1-1: General rules and rules for buildings. [31] CEN, European Committee for Standardization. EN 1990:2002. Eurocode 0: Basis of structural design.

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FIGURE CAPTIONS Fig. 1. Galvanized Ultra High Tensile Strength Steel (UHTSS) textile: cord detail (a), 12 cords/inch textile (S12, b), and 4 cords/inch textile (S4, c). Fig. 2. Testing setup for direct tensile tests on textile specimens (a,b) and composite specimens (c,d): overview of the setups (a,c) and detail of the clamping methods (b,d). Fig. 3. Stress-strain response curves of textile specimens under tension: unaged S12 and S4 (reference curves, a), S4 after artificial ageing in substitute ocean water for different durations (b), after 90° flexure (c), and after 90° flexure and ageing in substitute ocean water for 41 days (d). Fig. 4. Variation of tensile strength (a) and modulus of elasticity (b) of S4 textile specimens after flexure and/or ageing in substitute ocean water. Fig. 5. Scanning electron microscope (SEM) analysis of the steel cord after ageing in substitute ocean water for 41 days (a), flexure (b), and flexure with subsequent ageing (c). Fig. 6. Stress-strain response curves of composite specimens under tension: S12 textile with mineral mortar (S12M series, a) and mineral NHL mortar (S12L series, b), and S4 textile with mineral mortar (S4M series, c) and mineral NHL mortar (S4L series, d). Fig. 7. Crack pattern in composite specimens under tension: S12 textile with mineral mortar (S12M series, a) and mineral NHL mortar (S12L series, b), and S4 textile with mineral mortar (S4M series, c) and mineral NHL mortar (S4L series, d). Fig. 8. Testing setup for shear bond tests (a) and detail of the instruments for recording slip and strain (b). Fig. 9. Force-slip response curves of shear bond tests on reinforcement systems with S12 textile on modern strong brick (S12M-SB and S12L-SB series, a), modern weak brick (S12M-WB and S12L-WB series, b), historic brick (S12L-HB series, c), and tuff unit (S12M-TU and S12L-TU series, d). Fig. 10. Force-slip response curves of shear bond tests on reinforcement systems with S4 textile on modern strong brick (S4M-SB and S4L-SB series, a), modern weak brick (S4M-WB and S4L-WB series, b), historic brick (S4L-HB series, c), and tuff unit (S4M-TU and S4L-TU series, d). Fig. 11. Failure modes observed in shear bond tests: debonding with cohesive failure in the first layer of the substrate (mode A, S4M-WB series), debonding at the mortar-to-substrate interface (mode B, S4M-SB series), and shear failure at the textile-to-matrix interface (mode C, S4L-WB series). Fig. 12. Maximum force in shear bond tests for all specimens. Fig. 13. Specimen of S4M-TU series after failure and sketch of the stress field in the mortar: failure surface in the substrate (a) and in the detached SRG strip (b), cracks in the mortar matrix (c), and stresses in the mortar layer between textile and substrate (d). Fig. 14. Identification of engineering design parameters for SRG systems with S12 textile and mineral mortar (S12M, a), S12 textile and mineral NHL mortar (S12L, b), S4 textile and mineral mortar (S4M, c), and S4 textile and mineral NHL mortar (S4L, d).

TABLE CAPTIONS Table 1. Mechanical properties of mortar matrices and substrates (CV in round brackets). Table 2. Results of tensile tests on textile specimens (CV in round brackets). Table 3. Results of tensile tests on SRG specimens: stages I and II (CV in round brackets). Five tests were carried out for each series. Table 4. Results of tensile tests on SRG specimens: stage III and peak values (CV in round brackets). Five tests were carried out for each series. Table 5. Experimental plan for shear bond tests: acronyms of tested strengthening systems and considered substrates. Table 6. Results of shear bond tests (CV in round brackets). Five tests were carried out for each series. Table 7. Qualification parameters for tensile and shear bond failure.

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180

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(a) 160 140

140

120

120

100 80

100 80

60

60

40

40 Mineral mortar (S12M-SB)

20

Substrate: modern weak brick

160

Load [kN/m]

Load [kN/m]

(b)

Substrate: modern strong brick

Mineral mortar (S12M-WB)

20

Lime mortar (S12L-SB)

0 0

0.1

0.2 0.3 Slip [mm]

Lime mortar (S12L-WB)

0.4

0 0

0.5

180

0.1

0.2 0.3 Slip [mm]

0.4

180

(c)

(d) Substrate: tuff unit

160

Substrate: historic brick

140

140

120

120 Load [kN/m]

Load [kN/m]

160

100 80

100 80

60

60

40

40

20

Mineral mortar (S12M-TU)

20 Lime mortar (S12L-HB)

0 0

09/04/2015

0.5

0.1

0.2 0.3 Slip [mm]

0.4

Lime mortar (S12L-TU)

0.5

0 0

0.1

0.2 0.3 Slip [mm]

Risultati delaminazione S12

0.4

0.5

Figure 10

Paper SRG Tensile+Bond Figure 10

180

180

(a) 160 140

140

120

120

100 80

100 80

60

60

40

40 Mineral mortar (S4M-SB)

20

Substrate: modern weak brick

160

Load [kN/m]

Load [kN/m]

(b)

Substrate: modern strong brick

Mineral mortar (S4M-WB)

20

Lime mortar (S4L-SB)

0 0

0.5

1 Slip [mm]

Lime mortar (S4L-WB)

1.5

0 0

2

0.5

1 Slip [mm]

1.5

180

180

(c)

(d) 160

Substrate: historic brick

140

140

120

120 Load [kN/m]

Load [kN/m]

160

100 80

100 80

60

60

40

40

Substrate: tuff unit Mineral mortar (S4M-TU)

20

20 Lime mortar (S4L-HB)

0 0

09/04/2015

2

0.5

1 Slip [mm]

1.5

Lime mortar (S4L-TU)

2

0 0

0.5

Risultati delaminazione S4

1 Slip [mm]

1.5

2

Figure 11

Paper SRG Tensile+Bond Figure 11

Failure mode A

MSB-S4G

09/04/2015

Failure mode B

MSB-S4LS

Modalità di rottura

Failure mode C

MSB-S12LS

Figure 12

Paper SRG Tensile+Bond Figure 12

180 Failure mode A Failure mode B

Systems with S12

160

Systems with S4

Failure mode C

140

Load [kN/m]

120 100 80 60 40

Mineral mortar

Lime mortar

Modern Strong Brick

Mineral mortar

Lime mortar

Modern Weak Brick

Lime mortar Historic Brick

Mineral mortar

Lime mortar Tuff unit

20 0

09/04/2015 Figura

riepilogativa delle prove di bond (grafico a punti)

Figure 13

Struts

Cracks

Paper SRG Tensile+Bond Figure 13

Substrate

Mortar Textile

Crack

Crack

Tensile stress in the textile

(a)

09/04/2015

(b)

(c)

(d)

Figure 14

Paper SRG Tensile+Bond Figure 14

3500

3500 S12 with mineral mortar (S12M)

(a)

S12 with lime mortar (S12L)

(b)

3000

3000 stk

2500

stk etk

2

Stress [N/mm ]

etk

2000 500

1500

Modern weak brick Tuff

Stress [N/mm2]

400

1000

sdk

0.5

1500

400

1000

Modern strong brick

200

0 0

0 0

500

300

100

500

2000

Stress [N/mm2]

2

Stress [N/mm ]

2500

edk 0.05

0.1 0.15 Strain [%]

1 1.5 Strain [%]

2

300

Tuff

sdk

Historic brick Modern weak brick

100

500

0 0

0.2

0 0

2.5

Modern strong brick

200

0.5

edk 0.05

0.1 0.15 Strain [%]

1 1.5 Strain [%]

0.2

2

2.5

2

2.5

3500

3500

S4 with mineral mortar (S4M)

(c)

S4 with lime mortar (S4L)

(d) 3000

3000

stk

stk

2500 2

Stress [N/mm ]

2

Stress [N/mm ]

2500 2000

sdk Tuff

1500 Modern strong brick

2000 sdk

1500

Modern weak brick

Tuff

1000

1000

Modern strong brick

500 0 0

500

edk

0.5

etk

1 1.5 Strain [%]

etk

edk

2

2.5

Experimental envelope (Tensile tests on S12M and S4M) Experimental envelope (Tensile tests on S12L and S4L) Average curve (Tensile tests) Tensile failure

09/04/2015

Historic brick Modern weak brick

0 0

0.5

Shear Shear Shear Shear

failure failure failure failure

1 1.5 Strain [%] on on on on

modern strong brick modern weak brick historic brick tuff unit

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

Matrix / Substrate Matrix

Substrate

Mineral mortar Mineral-NHL mortar Modern strong brick Modern weak brick Historic brick Tuff unit

Acronym M L SB WB HB TU

fcm [N/mm2] 56.3 (2.7%) 20.6 (3.9%) 35.5 (12.9%) 14.7 (8.4%) 25.5 (11.4%) 4.4 (13.0%)

Ecm [kN/mm2] 22.01 (0.7%) 11.42 (5.0%) 3.45 (7.7%) 1.95 (12.0%) 3.21 (9.9%) 0.78 (10.7%)

ftm [N/mm2] 10.31 (2.6%) 5.42 (4.1%) 3.60 (11.7%) 1.88 (1.1%) 1.79 (3.7%) 0.56 (4.5%)

Table 1. Mechanical properties of mortar matrices and substrates (CV in round brackets).

Page 13 of 19

D [mm] 0÷0.5 0÷1.4 -

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

Specimen

Ageing

S12 S4 S4 S4 S4 S4 S4

None None S.O.W. 15 days S.O.W. 30 days S.O.W. 41 days 90° flexure 90° flexure + S.O.W. 41 days

Number of tests 6 6 3 3 3 3 3

ff [N/mm2] 3082.7 (1.6%) 3186.1 (0.8%) 3112.9 (0.9%) 3061.2 (0.4%) 3044.1 (1.0%) 2687.4 (3.1%) 2574.5 (1.0%)

Ff [kN/m] 789.2 267.6 261.5 257.1 255.7 225.7 216.3

εf [%] 2.26 (6.7%) 2.26 (3.6%) 2.11 (2.0%) 2.27 (4.9%) 2.20 (1.8%) 1.80 (3.1%) 1.82 (2.7%)

Table 2. Results of tensile tests on textile specimens (CV in round brackets).

Page 14 of 19

Ef [kN/mm2] 183.3 (2.2%) 183.7 (1.9%) 178.3 (0.8%) 176.8 (0.9%) 176.2 (0.6%) 165.7 (1.0%) 161.0 (0.2%)

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

SRG system S12M S12L S4M S4L

σI [N/mm2] 274.6 (3.9%) 89.3 (11.7%) 649.2 (11.5%) 121.1 (9.4%)

εI [%] 0.06 (9.8%) 0.04 (16.7%) 0.07 (36.4%) 0.05 (19.8%)

Et,I [kN/mm2] 629.7 (4.1%) 260.4 (16.0%) 1488.0 (7.9%) 292.6 (22.7%)

σII [N/mm2] 371.9 (18.8%) Undetectable 987.6 (6.8%) 205.5 (6.9%)

εII [%] 0.15 (34.0%) Undetectable 0.32 (14.3%) 0.07 (7.7%)

Et,II [kN/mm2] 159.9 (16.4%) Undetectable 120.3 (14.6%) 120.9 (16.0%)

Table 3. Results of tensile tests on SRG specimens: stages I and II (CV in round brackets). Five tests were carried out for each series.

Page 15 of 19

δ [mm] 24.4 (25.7%) 22.7 (11.1%) 22.0 (14.2%) 19.5 (27.0%)

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

SRG system S12M S12L S4M S4L

ft [N/mm2] 2925.9 (0.7%) 2852.6 (3.4%) 3021.8 (0.8%) 3242.8 (5.7%)

Ft [kN/m] 743.2 730.3 253.8 272.4

εt [%] 2.24 (1.5%) 2.21 (8.0%) 2.01 (3.9%) 2.17 (4.8%)

Et,III [kN/mm2] 162.7 (1.3%) 178.0 (2.8%) 172.4 (2.1%) 180.5 (1.9%)

Table 4. Results of tensile tests on SRG specimens: stage III and peak values (CV in round brackets). Five tests were carried out for each series.

Page 16 of 19

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

Steel textile 12 cords/inch 4 cords/inch

Mortar matrix

Strengthening system

Mineral Mineral HLM Mineral Mineral HLM

S12M S12L S4M S4L

Modern strong brick S12M-SB S12L-SB S4M-SB S4L-SB

Substrate Modern weak Historic brick brick S12M-WB − S12L-WB S12L-HB S4M-WB − S4L-WB S4L-HB

Tuff units S12M-TU S12L-TU S4M-TU S4L-TU

Table 5. Experimental plan for shear bond tests: acronyms of tested strengthening systems and considered substrates.

Page 17 of 19

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

Series

S12M-SB S12M-WB S12M-TU S12L-SB S12L-WB S12L-HB S12L-TU S4M-SB S4M-WB S4M-TU S4L-SB S4L-WB S4L-HB S4L-TU

Fd [kN/m] 98.0 (11.6%) 122.4 (5.9%) 123.9 (5.1%) 74.0 (8.9%) 77.2 (18.6%) 68.4 (8.3%) 113.8 (9.7%) 141.8 (5.4%) 114.7 (2.6%) 157.6 (3.6%) 106.3 (12.0%) 69.0 (10.0%) 111.7 (15.1%) 140.8 (11.3%)

sd [mm] 0.15 (29.2%) 0.15 (40.7%) 0.23 (16.2%) 0.10 (25.4%) 0.34 (19.7%) 0.16 (38.5%) 0.39 (15.0%) 1.27 (28.0%) 0.24 (52.0%) 0.82 (37.2%) 0.63 (23.2%) 0.47 (32.6%) 0.85 (32.8%) 1.76 (35.9%)

σd [N/mm2] 385.9 481.9 487.9 291.5 303.8 269.1 447.9 1688.3 1365.5 1876.3 1265.4 821.7 1329.5 1675.7

σd/ff [%] 12.5 15.6 15.8 9.5 9.9 8.7 14.5 53.0 42.9 58.9 37.9 25.8 41.7 52.6

Table 6. Results of shear bond tests (CV in round brackets). Five tests were carried out for each series.

Page 18 of 19

Failure mode C C C C-B C C C B A A-C C C B C

COST-D-15-00693 - DeSantis deFelice - MAIN DOCUMENT - REVISED.doc

Tensile failure εtk [%]

SRG system

σtk [N/mm2]

S12M

2746.9

1.87

697.7

S12L

2506.4

1.51

636.6

S4M

2834.2

1.65

238.1

S4L

2970.4

1.77

249.5

Ftk [kN/m]

Substrate

SB WB TU SB WB HB TU SB WB TU SB WB HB TU

Shear bond failure σdk εdk Fdk [kN/m] [N/mm2] [%] 281.4 0.07 71.5 415.8 0.17 105.6 430.5 0.18 109.3 231.0 0.11 58.7 172.3 0.08 43.8 217.1 0.11 55.1 346.9 0.18 88.1 1477.3 0.62 124.1 1284.4 0.50 107.9 1720.4 0.76 144.5 911.5 0.50 76.6 630.0 0.35 52.9 863.2 0.48 72.5 1233.4 0.68 103.6

Table 7. Qualification parameters for tensile and bond failure.

Page 19 of 19

Ed [kN/mm2] 402.0 244.6 239.2 210.0 215.4 197.4 192.7 238.3 256.9 226.4 182.3 180.0 179.8 181.4