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Article

Experimental Study and FEM Analysis on the Strengthening of Masonry Brick Walls Using Expanded Steel Plates and Shotcrete with and Without Glass Fiber Reinforcement

by
Zeynep Yaman
1,
Alper Cumhur
2,
Elif Ağcakoca
1,*,
Muhammet Zeki Özyurt
1,
Muhammed Maraşlı
3,
Mohammad Saber Sadid
1,
Abdulsalam Akrami
1 and
Azizullah Rasuly
1
1
Department of Civil Engineering, Sakarya University, 54200 Sakarya, Turkey
2
Department of Civil Engineering, Yalova University, 77200 Yalova, Turkey
3
Fibrobeton Company R&D Center, 81100 Duzce, Turkey
*
Author to whom correspondence should be addressed.
Buildings 2025, 15(15), 2781; https://doi.org/10.3390/buildings15152781
Submission received: 13 June 2025 / Revised: 23 July 2025 / Accepted: 25 July 2025 / Published: 6 August 2025
(This article belongs to the Special Issue Advanced Research on Concrete Materials in Construction)
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Abstract

In this study, an effective strengthening method was investigated to improve the seismic performance of masonry brick walls. The strengthening method comprised the use of shotcrete, which was applied in both glass fiber-reinforced and unreinforced forms for steel plates and tie rods. Thirteen wall specimens constructed with vertical perforated masonry block bricks were tested under diagonal compression in accordance with ASTM E519 (2010). Reinforcement plates with different thicknesses (1.5 mm, 2 mm, and 3 mm) were anchored using 6 mm diameter tie rods. A specially designed steel frame and an experimental loading program with controlled deformation increments were employed to simulate the effects of reinforced concrete beam frame system on walls under the effect of diagonal loads caused by seismic loads. In addition, numerical simulations were conducted using three-dimensional finite element models in Abaqus Explicit software to validate the experimental results. The findings demonstrated that increasing the number of tie rods enhanced the shear strength and overall behavior of the walls. Steel plates effectively absorbed tensile stresses and limited crack propagation, while the fiber reinforcement in the shotcrete further improved wall strength and ductility. Overall, the proposed strengthening techniques provided significant improvements in the seismic resistance and energy absorption capacity of masonry walls, offering practical and reliable solutions to enhance the safety and durability of existing masonry structures.

1. Introduction

Earthquakes are among the natural disasters that have serious effects on structures. During an earthquake, the strength of structures, energy absorption capacity, and ductility are important parameters that directly affect the safety of structures. One of the key structural parameters that contributes to the absorption and dissipation of earthquake energy is the lateral translational rigidity. Higher lateral rigidity limits excessive deformations during seismic loading and enabling more uniform energy dissipation throughout the structure. By restricting overall drift and inter-story displacements, this rigidity helps control the development and propagation of cracks, thereby reducing potential structural damage and enhancing overall seismic performance. Especially in masonry structures, strengthening the walls used as load-carrying elements provides significant contributions to the lateral translational rigidity of the structure, and the benefits of strengthening methods have been demonstrated in the literature. Masonry walls are structural elements that are generally produced from brittle materials and do not have a monolithic structure. Breakages and separations may occur due to the weak adhesion between the walls. Wall reinforcement applications contribute to solving this problem.
Various experimental studies have been conducted in the literature to increase the durability of masonry structures. Anania showed that the use of CFRP (Carbon Fiber-Reinforced Polymer) improves the outward behaviors in calcareous brick walls and that this material plays an important role in increasing the durability of the wall [1]. Calò studied the in-plane shear behavior of brick walls reinforced with IMC (Inorganic Mortar Composites) and stated that this type of reinforcement significantly increases the durability of the walls [2]. FRP reinforcements improve the reinforcement performance by increasing the shear capacity of the walls [3]. On the other hand, studies conducted to increase the shear capacity of the wall have led to the development of models to determine the carrying capacity of strengthened walls. Successful results were achieved in the evaluation of the shear capacity of the walls using simple mechanical models [4]. It was shown that the use of expanded steel sheets in brick walls provides significant benefits in terms of earthquake safety and emphasizes that such reinforcements enhance durability under horizontal loads [5,6,7]. Glass fiber-reinforced polymers significantly increase the carrying capacity of the wall in diagonal compression tests [8]. Today, the effectiveness of reinforcements made with different materials has been supported by various studies stating that they obtained higher strength in bonding tests performed on brick walls strengthened with FRP [9,10,11]. Diagonal compression tests on strengthened walls have been stated to provide a more sensitive approach for determining their load-carrying capacity [12]. The behavior of brick walls of different sizes under in-plane shear loads has been examined and it has been shown that the size of the wall can affect the reinforcement performance [13]. In addition, the performance of the walls in more detail has been evaluated, with in-plane and out-of-plane analyses using micro modeling methods [14]. It has been shown that reinforcement with composite materials is effective in increasing the strength of brick walls and that these reinforcement methods can be made more efficient [15,16]. While reference [15] focused on the role of composite materials in enhancing shear capacity, reference [16] demonstrated that such materials significantly improve the mechanical performance of walls in their experimental study on panels reinforced with basalt textile reinforced mortar. The role of bed-joint structures in the strengthening of brick walls has been investigated and the effects of such techniques on the shear capacity of the walls has been evaluated [17]. Many experimental studies have been carried out on the shear behavior of walls [18,19,20] stating that carbon fiber-reinforced polymers have an important place in strengthening brick walls and that these materials are effective in increasing durability. The diagonal compressive strength of single-headed masonry panels using mortar reinforced with industrial waste fibers has been investigated, and the authors of this study emphasized the critical effect of bed joints and coating reinforcements [21]. Another study evaluated the effects of coating and joint reinforcements on the carrying capacity of two-headed masonry walls strengthened with basalt and glass fiber-reinforced mortars [22]. These studies show that the use of innovative materials is effective in increasing structural durability. Studies such as [23] focused on modeling traditional material behavior by evaluating the diagonal tensile and shear strengths of rubble stone walls and mixed stone–brick masonry structures. The behavior of hollow clay brick panels under dynamic loads was investigated and the authors provided information on the energy absorption capacity of such structures [24,25]. Fiber reinforcement plays a critical role in increasing the durability of masonry panels. A few studies evaluated the performance of fiber-reinforced cementitious composite panels under seismic loads [26,27,28], while [29] investigated the effects of strengthening adobe structures with fiber reinforcement on ductility and strength. In addition, one study compared the effectiveness of different textile-reinforced mortar composites in increasing the shear capacity [30]. Numerical modeling studies also provide significant contributions to the understanding of strengthening techniques; [31,32] enveloped a damage-plasticity based interface model to model the in- and out-of-plane behavior of masonry panels, while [33] evaluated the ultimate shear capacity of perforated panels by inclined plane tests. A simplified micro-model to facilitate nonlinear static analysis of masonry panels was presented in another study [34]. These studies show that strengthening methods offer different approaches to improving the performance of masonry walls and that each method can be made more efficient in certain applications [35,36,37,38,39]. The diversity of strengthening techniques in improving the durability of masonry walls is demonstrated by the contributions of each method to different performance characteristics.
In this study, expanded steel sheets, one of the well-established techniques in the literature, were used to strengthen masonry walls. In the strengthening application, plate thickness and the spacing of tie rods, which anchor the plates to the wall, were defined as variable parameters. Additionally, instead of the commonly applied plaster layer in masonry structures, a 30 mm thick shotcrete layer was implemented, and to introduce a novel aspect, glass fiber reinforcement was incorporated into the shotcrete mix. The tensile strength of the tie rods, compressive strength of the masonry bricks, and mechanical properties of the mortar were evaluated through laboratory testing. The experimental setup was designed in accordance with ASTM E519 [40,41,42,43], and diagonal compression was applied using a steel frame system to simulate the effects of a reinforced concrete beam frame system on walls under the effect of diagonal loads caused by seismic loads. To further investigate the structural behavior of the test specimens, finite element models (FEM) were developed using ABAQUS software 2017. In these models, masonry bricks and mortar were characterized using elastic and plastic properties derived from experimental data, and both the expanded steel sheets and shotcrete layers were modeled using a micro-modeling approach. The validated FEM enabled the detailed analysis of load-carrying capacity, stiffness, ductility, and energy absorption capacity, along with critical parameters such as equivalent plastic strain (PEEQ), maximum principal stress, damage in plaster, and damage in bricks. By analyzing the behavior of reinforced walls, it aims to provide practical and low-cost solutions that will increase earthquake safety. The motivation of this study is to contribute to the development of effective strengthening strategies aimed at enhancing the seismic resistance of masonry structures under lateral loading conditions such as earthquakes (Figure 1).

2. Materials and Methods

Within the scope of this study, the expanded steel sheets reinforcement method, which is one of the accepted techniques in the literature, was used in the reinforcement of masonry walls built with block bricks, and in the experimental study, the plate thickness and the tie-rod spacing, which allows the boards to be mounted on the wall, were determined as variable parameters. In addition, shotcrete application, which is preferred in the reinforcement of masonry buildings, was also used in this study. In order to add a more original value to the work, reinforcement was made by adding glass fibers to shotcrete. The specimens were produced in the laboratory of the Sakarya University Civil Engineering Department. The walls were formed with 13.5 × 19 × 29 cm vertical perforated masonry block bricks with dimensions of 1000 × 1000 mm. Expanded steel sheets in 3 different thicknesses with thicknesses of 1.5 mm, 2 mm, and 3 mm with gaps of 54 × 26 mm were used as reinforcement plates. The anchoring of the boards to the wall was provided using tie-rod shafts with a diameter of 6 mm. The names and descriptions of the experimental specimens used in the study are given in Table 1. The characteristic properties of the materials used in the experimental study were examined by laboratory experiments.

2.1. Expanded Steel Sheets

Expanded steel sheets are materials manufactured by drilling stainless steel sheets with a raised or flattened open diamond pattern. Steel plates with raised wavy ridges for high adherence were used in the study. The steel plate used in the experimental study and its dimensions are given in Figure 2. In order to test the tensile strength of the steel plate used, samples were cut in accordance with ASTM E8:2016 [41] standards and a tensile test was performed.

2.2. Tie Rod

For anchoring the steel plates to the wall samples, tie rod with a diameter of 6 mm was used. In order to ensure the healthy anchoring process and to ensure adherence with the shotcrete applied to the wall surfaces, tie-rod lengths were determined as 250 mm by calculating 15 mm end protrusions from the outer surface of the plate after the anchoring process. The number of tie rods specified in the test matrix was used for each wall test specimen. A tensile test was performed to determine the strength of tie rods and the results are presented in Table 2.

2.3. Brick

Within the scope of the experimental study, 135 × 190 × 290 mm vertical perforated masonry block bricks were used in the wall samples. In order to test the properties of vertical perforated masonry bricks according to EN 1052-1 [44], pressure tests were carried out on 27 bricks. The weight of the bricks was measured as approximately 4.75 kg/piece on average, and the compressive strength test results are given in Table 3.

2.4. Mortar

In the study, a mortar obtained by mixing slaked lime and cement with water and sand was used. The mortar used in the test samples was prepared in a way that the weight ratios were 20% cement, 60% sand, 10% lime, and water. Nine samples were obtained from each of the mortars prepared at three different times in accordance with EN 998-2 [45,46]. After waiting for 5 days in a mold in a 95–100% humidity curing environment and 23 days in a 65–70% humidity environment, it was subjected to pressure testing. The test results were analyzed by taking the mean, standard deviation, and variance values (Table 4).

2.5. Shotcrete

In this study, 30 mm thick wet shotcrete was applied to the prepared reference and reinforced samples instead of plaster. Shotcrete application was carried out by FİBROBETON INC.
The shotcrete mixture used in the study contains 36% cement, 4% meta kaolin, 40% silica sand, 2.12% hyper plasticizer, 1.2% acrylic polymer, and 12.8% water. For glass fiber shotcrete, which is one of the original values of the study, 3.88% ARG glass fiber was added to this mixture. In shotcrete, CEM II/B-L 42.5R White Portland Calcareous Cement compatible with EN 197-1 [46], silica sand as aggregate suitable for GRC-CTB production, metakaolin (MTK), and polycarboxylate ether-based plasticizer additive were used to increase the durability, waterproofing, and chemical resistance of the concrete. The alkali-resistant (AR) glass fibers used belong to NEG company (https://www.neg.co.jp/en/product/fiber/a-chopped_strand, accessed on 1 June 2025) and were supplied by Fibrobeton INC. In the study, 4-cylinder samples of ARG fiber and fiber-free sprayed concrete were subjected to 28-day compressive strength testing in accordance with EN 206-1 [47] and the test results are given in Table 5. Additionally, bending tests were performed on test specimens made of fibrous and non-fibrous concrete according to the same standard.

3. Experimental Setup

Diagonal pressure test methods on walls are given in ASTM E519 [40,41,42,43]. In this method, the upper and lower corners of the specimen are limited and the load is transmitted to the specimen by means of loading heads, so that the wall side surfaces are free. In the actual building behavior, the pressure contact surface between the masonry walls surrounded by vertical and horizontal beams and these reinforced concrete structural specimens are formed under the effect of horizontal loads. Since the pressure contact surface may enlarge due to deformation during an earthquake, this situation was simulated in the experimental setup and a square steel frame with four sides articulated and rigid surfaces was used to transfer the diagonal component of the horizontal earthquake load to the specimen [5].
The joints at the four corners of the steel frame ensured that the entire applied diagonal load was transferred directly to the specimen and that the steel frame did not resist the load. The experimental mechanism is shown in Figure 3. There are a total of four steel legs, which are expressed as pendulum feet, in the experimental setup. The upper and lower stiffening rods in the experimental setup can be fixed by means of these steel legs.
The static charge was given to the specimens by means of a movable double-acting hydraulic piston located at the bottom of the experimental mechanism with a capacity of 1000 kN. In the experimental mechanism, a total of 4 LVDTs, each with a precision of 0.01 mm, were used to measure the shortening of the specimen in the direction of the load and the elongation of the surface perpendicular to the load (horizontal).
During the experiment, electrical signals from these devices were digitized and transferred to the computer using the TestBox1001 Series data acquisition system (by TDG), which offers 16-bit resolution, high measurement performance, wide sensor compatibility, direct sensor connection capability, and supports 8 × 8 channels with single-run operation and connection to all channels via a single USB port (Figure 4). To measure the vertical displacements at the lower corners of the stiff plate mounted on the hydraulic jack, two LVDTs with a measurement capacity of 400 mm were used. Additionally, two LVDTs with a capacity of 50 mm were fixed on the front and rear surfaces of the specimen to measure the lateral deformations perpendicular to the loading direction (Figure 4). In the experimental loading program, at the beginning of the experiment, the loading up to 10 mm deformation was carried out with a 5% increase rate, the loading speed of 10–35 mm deformation was carried out with a 25% increase rate, and after the deformation exceeded 35 mm, the loading was carried out with a constant 50% increase rate.

4. Experimental Observations and Discussion

Within the scope of the study, 13 specimens composed of vertical perforated block bricks, expanded steel sheets, and shotcrete were subjected to static diagonal loading tests. Shotcrete and fiber-shotcrete-reinforced specimens were compared with plastered reference specimen. In addition, tie-rod spacing and expanded steel sheet thickness were determined as variable parameters in the reinforced specimens. Load–displacement graphs of all specimens are given in Figure 5.
MW, MWS and MWSG specimens were created to see the effects of shotcrete and glass fiber shotcrete application on wall behavior. The damages and behaviors observed during the experiment in all specimens are summarized below.
All experimental specimens were created to see the effects of shotcrete and glass fiber shotcrete application on wall behavior. The crack, damage, and behaviors observed during the experiment in all specimens are summarized below.
In the MW specimen, during the experiment, various types of damage were observed, accompanied by increased load and displacement. At a load of 60 kN and a displacement of 31 mm, fine hairline cracks occurred in the gypsum plaster. With a load of 85 kN and a displacement of 36 mm, the first main crack was formed 20 cm to the right of the diagonal parallel to the loading axis. The width of this crack was about 3 mm. At a load of 86 kN and a displacement of 38 mm, swellings were observed in the plaster at the lower corner of the wall, and these swellings became more pronounced as the load increased. With a load of 98 kN and a displacement of 38 mm, a second large crack occurred on the rear surface of the wall, 15 cm to the right of the diagonal parallel to the loading axis, and the crack width was recorded as 3 cm. At this point, the breaking sounds of the bricks were heard and the load-carrying capacity decreased. At a load of 115 kN and a displacement of 25 mm, the wall completely lost its strength and the experiment was terminated.
In the MWS specimen, with a load of 103 kN and a displacement of 14 mm, a vertical crack appeared on the rear surface 30 cm to the left of the loading axis, resulting in a 3 kN reduction in payload. At a load of 70 kN and a displacement of 21 mm, a crack was observed on the frontal surface 50 cm below the load impact zone. Three hairline cracks were detected on the posterior surface with a load of 66 kN and a displacement of 22 mm. At a load of 60 kN and a displacement of 50 mm, the crack in the rear surface became apparent, and at a load of 66 kN and a displacement of 27 mm, a horizontal crack grew rapidly and there was a sudden loss of load. With a load of 55 kN and a displacement of 66 mm, shotcrete layers began to separate on the front surface, and the crack on the rear surface expanded, reducing the load-carrying capacity.
In the MWSG specimen, at the beginning of the experiment, at a load of 125 kN and a displacement of 20 mm, slight spillage was observed in the lower and upper joints. At a load of 150 kN and a displacement of 25 mm, a sudden loss of charge of about 5 kN and crushing and fractures occurred in the bricks. At a load of 165 kN and a displacement of 75 mm, there were fractures in the bricks, but no cracking in the shotcrete layer; crushing was detected only in the joints. At a load of 73 kN and a displacement of 147 mm, a crack appeared in the front face upper joint, and the bricks were not separated from the spray layer. At a load of 58 kN and a displacement of 175 mm, swelling, fracture, and crushing were observed. Although the bricks were broken, there was no damage to the shotcrete layer. The experiment was terminated when a load of 60 kN and a displacement of 240 mm were reached.
In the MWS 2-200 specimen, with a displacement of 20 mm and a load of 177 kN, a small hairline crack occurred on the back surface of the wall. With a displacement of 26 mm and a load of 199 kN, the shotcrete swelled at the front surface upper joint. There was a small load loss at a displacement of 35 mm and a load of 273 kN, with a sudden load loss of about 30 kN at a displacement of 40 mm. In the meantime, crushing of the joints and fractures in the shotcrete occurred. At the end of the experiment, large crushing and fractures were observed in the upper joints of the wall, small crushing and fractures were observed in the lower joints, and crumpling was observed in the steel plate. There was no stripping or rupture of the anchor bolts. In the areas where the shotcrete came into contact with the steel frame, the particles fell, while there were no cracks or breakages on the middle surfaces.
Figure 6 presents the sequential visuals illustrating the crack initiation, propagation, and overall damage status of the MW, MWS, MWSG, and MWS 2-200 specimens.
In the MWS 3-400 specimen, at a displacement of 18 mm and under a load of 178 kN, crackling noises were heard in the masonry due to stress build-up. The load increased to 214 kN at a displacement of 25 mm, cracking sounds continued to be heard, and a small loss of load was observed. At a displacement of 26 mm, a small fracture occurred in the shotcrete due to crushing, resulting in a sudden loss of load. As the load gradually decreased, it was observed that the shotcrete layers separated, the steel plate crumpled, and there were fractures in the vertical perforated bricks. At a displacement of 105 mm, the load decreased to 100 kN, but under the influence of steel plates and bolts, the load increased again and reached 122 kN at a displacement of 170 mm. At a displacement of 240 mm, the load was measured as 88 kN and the experiment was terminated due to the end of the stroke length of the piston. While there was no serious damage in the load-carrying areas, a large fracture of the shotcrete was observed in the lower joint.
In the MWS 3-200 specimen, at a displacement of 21 mm and under a load of 174 kN, small crushes and swellings were observed in the shotcrete at the rear surface lower joint. The load increased to 206 kN at a displacement of 30 mm, and shotcrete swelled and poured at the rear surface upper joint. At a displacement of 32 mm and a load of 201 kN, the front surface had blisters on the lower joint, wrinkles were observed on the steel plate, but there was no stripping on the bolts. When the load reached 213 kN, crushing of the masonry and fractures in layers of shotcrete continued. At a displacement of 45 mm, the load was reduced to 197 kN, and at a displacement of 240 mm, the experiment was terminated due to the stroke length of the piston at a load of 153 kN.
In the MWS 3-150 specimen, at a displacement of 25 mm and a load of 200 kN, minor crushes occurred in the wall joints, but there was no breakage or fragmentation. At a displacement of 26 mm, the load decreased from 203 kN to 197 kN due to shotcrete swelling at the upper joint, and then the wall took the load and reached 222 kN at 40 mm displacement. At this time, swelling of the lower and upper joints and pouring of small pieces of concrete were observed. At a displacement of 58 mm, when the load was around 206 kN, there was a small loss of load due to the fall of the concrete in the rear upper joint, wrinkling of the steel plate, but no stripping of the bolts. The load gradually decreased to 152 kN, while damage to the wall increased. In general, there were no sudden large load losses, and the load increased again with the increase in contact surfaces and the introduction of anchor bolts. The experiment was terminated at a displacement of 240 mm and it was determined that the wall carried a load of 205 kN.
In the MBWS 3.0-100 specimen, at a displacement of 35 mm and a load of 243 kN, a small fracture occurred on the rear surface of the wall. At a displacement of 50 mm, under a load of 266 kN, the shotcrete on the front surface of the wall fractured, but despite the sudden load losses, the wall continued to carry a load of up to 274 kN at a displacement of 55 mm. In the 55–60 mm displacement range, a load loss of approximately 30 kN was experienced due to breakage of shotcrete or wrinkling of steel plates, reducing to 244 kN. While the damages and the compression of the wall continued, the load gradually increased with the introduction of steel plates and fasteners, and since the experiment was terminated due to the limited stroke length at a displacement of 240 mm, it was determined that the wall carried a load of 244 kN.
Figure 7 presents the sequential visuals illustrating the crack initiation, propagation, and overall damage status of the MWS 3-400, MWS 3-200, MWS 3-150, and MWS 3-100 specimens.
In the MWSG 1.5-150 specimen, at a load of 120 kN and a displacement of 15 mm, no crushing was observed on the back and front faces, lower and upper joints, and the specimen maintained its integrity. At a load of 162 kN and a displacement of 22 mm, crushes were observed in the front face lower joint due to compression, and a horizontal fracture of approximately 3 cm was observed in the upper joint. On the back surface, swelling formed on the lower and upper joints as a result of crushing the spray layer, but the wall integrity was preserved. At a load of 150 kN and a displacement of 123 mm, no cracking, rupture, or slippage was observed on the front and rear surfaces, and the specimen maintained its integrity.
In the MWSG 2-150 specimen, at a load of 127 kN and a displacement of 20 mm, no crushing was observed in the lower and upper joints and the specimen maintained its integrity. At a load of 238 kN and a displacement of 53 mm, swelling due to crushing was observed in the lower and upper joints of the front face, crushing of the lower joint of the rear face, and 1.5 cm breakage in the upper joint. At a load of 184 kN and a displacement of 138 mm, it was observed that the shotcrete was separated from the reinforcement plate by 2 cm in the lower joint of the front face, and swelling occurred in the lower and upper joints of the back face. At a load of 165 kN and a displacement of 205 mm, no cracking was observed in the specimen.
In the MWSG 2-200 specimen, at a load of 117 kN and a displacement of 20 mm, no crushing was observed in the lower and upper joints and the specimen maintained its integrity. At a load of 180 kN and a displacement of 28 mm, crushing occurred in the front face lower joint, but no cracks occurred. At a load of 254 kN and a displacement of 65 mm, a 5 cm break in the upper joint, wrinkling of the steel plate, and separation from the wall were observed on the rear surface. At a load of 152 kN and a displacement of 191 mm, the front face lower joint swelled. The maximum load is recorded as 255 kN and the maximum deformation as 60 mm.
In the MWSG 3-150 specimen, no crushing was observed at a load of 160 kN and a displacement of 20 mm. At a load of 190 kN and a displacement of 23 mm, slight crushing occurred on the front and rear surfaces. At a load of 260 kN and a displacement of 40 mm, crushes occurred on both surfaces and a sudden loss of load was observed. At a load of 222 kN and a displacement of 103 mm, a 2 cm fracture occurred in the upper joint of the front face, but no cracks were seen in the middle part. At a load of 202 kN and a displacement of 140 mm, a block-like separation was observed at the rear surface upper joint. The maximum load is recorded as 270 kN, the deformation is 50 mm.
Figure 8 presents the sequential visuals illustrating the crack initiation, propagation, and overall damage status of the MWSG 1.5-150, MWSG 2-150, MWSG 2-200, and MWS 3-150 specimens.
In the MWSG 3-400 specimen, at a load of 105 kN and a displacement of 20 mm, slight spillage was observed in the lower and upper joint. At a load of 144 kN and a displacement of 30 mm, an abrupt load loss of about 10 kN occurred, but no cracking or blistering was observed. At a load of 170 kN and a displacement of 64 mm, heaving was observed in the right joint due to insufficient adherence between the glass fiber-reinforced shotcrete and the steel plate. At a load of 120 kN and a displacement of 85 mm, the spray layer was separated from the steel plate, but no cracking or blistering occurred. At a load of 102 kN and a displacement of 95 mm, there was a sudden loss of load due to loss of adherence, crushing, and swelling of the front and rear joints. At a load of 84 kN and a displacement of 140 mm, crushing and swelling occurred on the front and rear surfaces, but the specimen lost its integrity due to loss of adherence. Figure 9 presents the sequential visuals illustrating the crack initiation, propagation, and overall damage status of the MWS 3-400 specimen.
At the end of the experimental study, the load-carrying capacity, stiffness, ductility, and energy conversion capacities of all experimental specimens are given in Table 6. When MWS, MWS 3-100, MWS 3-150, MWS 3-200, and MWS 3-400 specimens are examined, it is observed that the use of tie rod causes an increase of 58% in MWS 3-400, 110% in MWS 3-200, 120% in MWS 3-150, and 190% in MWS 3-100 in carrying capacity compared to the MWS specimen. In stiffness compared to the MWS specimen, there is an increase of 48% in MWS 3-400, 88% in MWS 3-200, 98% in MWS 3-150, and 108% in MWS 3-100, and in ductility, it was observed that there was a 20% increase in MWS 3-400, 42% in MWS 3-200, 49% in MWS 3-150, and 82% in MWS 3-100. Compared to the MWS specimen, it was observed that the energy absorption capacities increased by 1.06 times in MWS 3-400, 1.75 times in MWS 3-200, 1.84 times in MWS 3-150, and 2.71 times in MWS 3-100. A similar situation was observed in MWSG 2-200 and MWSG 2-150 specimens. When the MWS 2-200 and MWSG 2-200 test specimens were compared, it was observed that there was an approximate 16% increase in carrying capacity, a 20% decrease in stiffness, a 2% increase in ductility value, and a 15% increase in energy absorption capacity. When comparing MWS 3-150 and MWSG 3-150 test specimens, it is observed that there is an approximate 17% increase in carrying capacity, 27% decrease in rigidity, 3% increase in ductility value, and 18% increase in energy absorption capacity. When comparing MWS 3-400 and MWSG 3-400 test specimens, it is observed that there is an approximate 25% increase in carrying capacity, 26% decrease in rigidity, 33% increase in ductility value, and 5% increase in energy absorption capacity.
An examination of the effect of plate thickness clearly demonstrates that increasing the thickness substantially enhances the load-carrying capacity, ductility, and energy absorption capacity of masonry walls. In the MWS specimen reinforced with steel plates, when the tie-rod spacing was held constant and the plate thickness was increased by 50%, improvements of approximately 7% in load-carrying capacity, 19% in stiffness, 28% in ductility, and 5% in energy absorption capacity were achieved. Similarly, in the MWSG specimen reinforced with both steel plates and fibers, a 50% increase in plate thickness led to increases of approximately 9% in load-carrying capacity, 11% in stiffness, 4% in ductility, and 9% in energy absorption capacity. Furthermore, thicker plates contributed significantly to enhancing the overall stability and ductile behavior of the walls, thereby improving their energy dissipation capabilities under static diagonal loading.
The influence of tie-rod spacing emerged as another critical parameter governing wall performance. Reducing the tie-rod spacing markedly improved the structural characteristics, including load-carrying capacity, stiffness, ductility, and energy absorption. In the steel plate-reinforced MWS specimen, when the plate thickness was kept constant and the tie-rod spacing was increased by 50%, decreases of approximately 30% in load-carrying capacity, 5% in stiffness, 22% in ductility, and 30% in energy absorption capacity were observed. When the spacing was increased fourfold, these reductions became even more pronounced, reaching approximately 83% in load-carrying capacity, 41% in stiffness, 52% in ductility, and 80% in energy absorption capacity. In the MWSG specimen with steel plate and fiber reinforcement, increasing the tie-rod spacing by 2.7 times resulted in decreases of approximately 30% in load-carrying capacity, 31% in stiffness, and 57% in energy absorption capacity, with no significant change observed in ductility.
These findings clearly highlight that a denser tie-rod arrangement provides more effective confinement, mitigates stress concentrations, and significantly improves the overall integrity and seismic performance of masonry walls. Therefore, the optimization of both plate thickness and tie-rod spacing is essential for achieving superior structural behavior and enhanced to static diagonal loading.

5. Finite Element Modeling Analysis

In this part of the study, the structural behavior of masonry walls strengthened with expanded steel plates and shotcrete was analyzed using the finite element method (FEM) in ABAQUS. A total of 13 three-dimensional (3D) FEM models were developed to reproduce the experimental setup. The CDP (Concrete Damaged Plasticity) property of ABAQUS was used to model the masonry bricks and plaster, accurately capturing their nonlinear behavior under loading. Despite its origins in concrete modeling, the CDP model is well suited to masonry because it can capture nonlinear phenomena such as tensile cracking, compressive crushing, and stiffness degradation—key features of masonry behavior under cyclic and dynamic loads. Several researchers have effectively used CDP in masonry modeling, taking advantage of its ability to represent localized cracking, damage accumulation, and plastic deformation in both brick-and-mortar materials [48,49]. The expanded steel plates were modeled with varying thicknesses of 1.5 mm, 2.0 mm, and 3.0 mm. A micro-modeling approach was employed, where the mortar between bricks was explicitly modeled using cohesive properties to capture localized failures, such as cracking and separation. Micro-modeling involves explicitly representing individual masonry units and mortar joints, allowing detailed simulation of interface behavior, localized cracking, and separation. This approach is computationally intensive but offers high fidelity in capturing local damage mechanisms. In contrast, macro-modeling treats masonry as a homogeneous or smeared continuum with equivalent material properties. This simplification reduces computational demands and is often more practical for large or complex structures, especially for global seismic assessment and vulnerability studies. Macro-modeling is widely used for territorial-scale analyses of historical buildings and churches, where the main goal is to assess overall stability and collapse mechanisms rather than local crack patterns [50,51]. Steel sheets expanded with diamond-shaped holes to increase the interaction of steel sheets with shotcrete and wall (Figure 10).
All material properties used in the FEM models were derived from experimental tests conducted as part of this research. Masonry properties, such as Young’s modulus, compressive strength, and tensile strength, were obtained through tests on vertically perforated bricks (Table 3). Expanded steel plates were characterized using dog-bone tensile tests, providing yield stress, tensile strength, and modulus of elasticity values (Figure 2).
For the shotcrete, both glass fiber-reinforced and non-reinforced variants were tested for compressive strength, density, and elastic modulus (Table 5). The data obtained from these tests were used in the FEM modeling, as detailed in the experimental study section of this article.
The boundary conditions were designed to mimic the experimental setup [52,53]. The steel lateral brace constrained the masonry walls, with its base fixed to provide stability. Additionally, all corners of the brace were modeled as hinges to allow rotational freedom, accurately replicating the experimental setup (Figure 10). This configuration ensured the realistic transfer of diagonal compression forces and associated deformations in the FEM model (Figure 11, Figure 12 and Figure 13).
The models were subjected to displacement-controlled static loading, with a total diagonal displacement of 250 mm applied at a 45° angle. The interaction between various components was critical to accurately simulating the system’s response. Cohesive properties defined the mortar between bricks and the interface between plaster and masonry. A hard contact interaction with a friction coefficient of 0.4 was specified between the steel brace and the masonry wall to prevent unrealistic penetration or slipping. Tie constraints were applied to represent the anchors connecting the steel plates to the walls, with the number of ties varying based on the model configuration.
Mesh sensitivity studies ensured the accuracy of the FEM models while maintaining computational efficiency. A uniform mesh size of 20 × 20 mm was selected, providing a balance between computational cost and result precision. The masonry walls and plaster layers were discretized using C3D8R elements, while T3D2 elements were used for the expanded steel plates. The FEM results, including stress–strain distributions, load–displacement behavior, and failure patterns, were compared with experimental data. The FEM models closely matched the experimental results, with discrepancies limited to 5–10%, confirming the validity of the numerical approach. Stress concentrations were observed around the ties and steel plates, indicating their critical role in enhancing the wall’s load-carrying capacity. Displacement patterns revealed the effectiveness of shotcrete and steel plates in mitigating deformation. Failure modes included cracking in the masonry walls and separation at the mortar interfaces, closely replicating experimental observations. The damage distribution results shown in Figure 11, Figure 12 and Figure 13 highlight compressive crushing in brick units, cracking and separation at mortar interfaces, and localized yielding of steel plates. Stress concentrations are observed near the tie anchor points and around the perforations of the expanded steel plates, indicating their critical role in force transfer and failure initiation. These simulated damage patterns are consistent with experimental observations, replicating the initiation and progression of cracking, and providing insight into the failure mechanisms of the strengthened masonry walls.
Comparative load–displacement graphs for FEM and experimental results, along with stress distribution plots, are presented in Figure 14 and Figure 15, illustrating the accuracy and reliability of the FEM approach in capturing the complex interactions within the reinforced masonry system. The comparative in Figure 14 and Figure 15 demonstrate overall good agreement between FEM predictions and experimental results, with discrepancies generally within 5–10%. These differences can be attributed to several factors, including material property heterogeneity not fully captured in input parameters, the approximation of discrete crack paths by smeared or cohesive models, simplifications in contact interaction definitions, and idealized boundary conditions. Additionally, the use of a uniform 20 mm mesh may limit the resolution of local failure mechanisms, while the tie constraints in the model assume perfect anchorage without local yielding or slip. Despite these limitations, the FEM results effectively reproduce the global structural response and validate the modeling approach. The convergence rates of the experimental and FEM models in Pmax values are given in Table 7.

6. Conclusions

The results obtained from the diagonal axial compression tests and finite element model analysis of 13 masonry brick walls manufactured with 4 different variables, namely sprayed concrete, glass fiber, tie-rod spacing, and steel plate thickness, are summarized below.
  • In comparisons among MW, MWS, and MWSG specimens, the MWS specimen showed a 30% increase in load-carrying capacity, while the MWSG specimen achieved a 70% increase compared to MW. Additionally, rigidity improved by 60% in MWS and 90% in MWSG; ductility increased by 20% and 40%, respectively; and energy absorption capacity rose by 40% and approximately 1.5 times. These findings clearly highlight the significant positive effects of sprayed concrete and glass fiber-reinforced sprayed concrete on the mechanical performance of masonry walls.
  • Comparison of MWS 2-200, MWS 3-150, and MWS 3-400 test specimens with the MWSG series specimens revealed that glass fiber material incorporated into shotcrete significantly improved the performance of masonry walls. Load-bearing capacity increased by 16% to 25%, while stiffness decreased by 20% to 27%. Ductility increased by 2% to 33%, and energy dissipation capacity increased significantly by 15% to 33%.
  • When MWS and MWS 3-100, MWS 3-150, MWS 3-200, and MWS 3-400 specimens were examined, it was observed that the use of tie rods had positive effects on the carrying capacity, rigidity, ductility, and energy absorption capacity of masonry walls. With the addition of tie rods, an increase from 58% to 190% was achieved in the carrying capacity, and this increase rate increased significantly with the decrease in the tie-rod spacing. Similarly, an increase from 48% to 108% was recorded in rigidity and from 20% to 82% in ductility. In addition, increasing the tie-rod frequency increased the energy absorption capacity from 1.06 times to 2.71 times.
  • When comparing MWSG 1.5-150 and MWSG 3-150 wall specimens, increasing slab thickness results in approximately a 22% increase in bearing capacity, an 18% increase in stiffness, a 10% increase in ductility, and a 24% increase in energy absorption capacity.
  • The MW specimen exhibited brittle behavior after the vertical main crack during the test, exhibiting sudden strength loss and limited displacement. In contrast, all of the strengthened walls exhibited a ductile behavior, maintaining their integrity up to maximum displacement and maintaining their load-carrying capacity without experiencing sudden load loss.
  • It has been observed that the addition of glass fiber to shotcrete causes an increase in the carrying capacity of masonry brick walls by 16–25%, an increase in their ductility by 2–33%, an increase in their energy absorption capacity by 15–33%, and a decrease in their stiffness by 20–27%.
  • Comparison of experimental (EXP) and finite element method (FEM) results showed small differences in load capacity, ranging from 0.2% (MW) to 8.9% (MWS). These findings confirm that FEM accurately predicts the load capacity of masonry walls and aligns well with experimental data, proving to be a reliable and effective evaluation method.
Recommendations: Specifically, it is suggested to explore different configurations of tie-rod spacing and plate thickness ratios in larger-scale masonry wall applications, to investigate alternative fiber types and shotcrete compositions for further performance enhancement, and to perform long-term durability studies under cyclic and environmental loading. These suggestions aim to guide future studies in optimizing the seismic strengthening of masonry structures.

Author Contributions

Z.Y.: writing—review and editing, writing—original draft, visualization, supervision, investigation, data curation, conceptualization, methodology, experimental procedure. A.C.: writing—review and editing, writing—original draft, visualization, supervision, investigation, data curation, conceptualization, methodology. M.Z.Ö.: carrying out the tests, experimental procedure. E.A.: writing—review and editing, writing—original draft, visualization, supervision, investigation, data curation, conceptualization, methodology, experimental procedure. M.S.S.: Writing—review and editing, writing—software, investigation. M.M.: material procurement, preparation of test specimens, conducting the experiments. A.A.: carrying out the tests, experimental procedure. A.R.: carrying out the tests, experimental procedure. All authors have read and agreed to the published version of the manuscript.

Funding

This research received no external funding. The APC was funded by Sakarya University.

Data Availability Statement

The data that support the findings of this study are available from the corresponding author upon reasonable request.

Acknowledgments

The authors would like to thank Fibrobeton Inc. (https://fibrobeton.com/) for their valuable support in providing fiber-reinforced concrete.

Conflicts of Interest

Author Muhammed Maraşlı was employed by the company Fibrobeton Company R&D Center. The remaining authors declare that the research was conducted in the absence of any commercial or financial relationships that could be construed as a potential conflict of interest.

References

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Figure 1. Flow chart of the study.
Figure 1. Flow chart of the study.
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Figure 2. Mechanical properties of plate perforations.
Figure 2. Mechanical properties of plate perforations.
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Figure 3. Experimental setup and a sample test specimen.
Figure 3. Experimental setup and a sample test specimen.
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Figure 4. Measuring instruments.
Figure 4. Measuring instruments.
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Figure 5. Load–displacement graphs obtained from the experiments.
Figure 5. Load–displacement graphs obtained from the experiments.
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Figure 6. Damage status of MW (a–c), MWS(d–f), MWSG (g–i), and MWS 2-200 (j–l), specimens.
Figure 6. Damage status of MW (a–c), MWS(d–f), MWSG (g–i), and MWS 2-200 (j–l), specimens.
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Figure 7. Damage status of MWS 3-400 (a–c), MWS 3-200 (d–f), MWS 3-150 (g–i), and MWS 3-100 (j–l) specimens.
Figure 7. Damage status of MWS 3-400 (a–c), MWS 3-200 (d–f), MWS 3-150 (g–i), and MWS 3-100 (j–l) specimens.
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Figure 8. Damage status of MWSG 1.5-150 (a–c), MWSG 2-150 (d–f), MWSG 2-200 (g–i), and MWSG 3-150 (j–l) specimens.
Figure 8. Damage status of MWSG 1.5-150 (a–c), MWSG 2-150 (d–f), MWSG 2-200 (g–i), and MWSG 3-150 (j–l) specimens.
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Figure 9. Damage status of MWSG 3-400 (a–c), specimens.
Figure 9. Damage status of MWSG 3-400 (a–c), specimens.
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Figure 10. FEM configuration of strengthened masonry walls.
Figure 10. FEM configuration of strengthened masonry walls.
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Figure 11. Damage results of finite element model analysis of MW, MWS, and MWSG reference test specimens.
Figure 11. Damage results of finite element model analysis of MW, MWS, and MWSG reference test specimens.
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Figure 12. Damage results of finite element model analysis of all MWS test specimens.
Figure 12. Damage results of finite element model analysis of all MWS test specimens.
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Figure 13. Damage results of finite element model analysis of all MWSG test specimens.
Figure 13. Damage results of finite element model analysis of all MWSG test specimens.
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Figure 14. Comparative load–displacement graphs for FEM and experimental results of MW, MWS, and MWSG reference test specimens.
Figure 14. Comparative load–displacement graphs for FEM and experimental results of MW, MWS, and MWSG reference test specimens.
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Figure 15. Comparative FEM and experimental results of reinforced test specimens.
Figure 15. Comparative FEM and experimental results of reinforced test specimens.
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Table 1. Specimen configuration and parameters for masonry wall samples and labeling.
Table 1. Specimen configuration and parameters for masonry wall samples and labeling.
Buildings 15 02781 i001SpecimenPlate Thickness
(mm)		Anchor
Spacing
(mm)		Number of
Tie Rods
MW---
MWS---
MWSG---
MWS 2-2002.020025
MWS 3-4003.04009
MWS 3-2003.020025
MWS 3-1503.015049
MWS 3-1003.0100100
MWSG 1.5-150 1.5 150 49
MWSG 2-1502.015049
MWSG 2-2002.020025
MWSG 3-1503.015049
MWSG 3-4003.04009
Table 2. Axial tensile test results for tie rods.
Table 2. Axial tensile test results for tie rods.
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Buildings 15 02781 i003Max Force
(kN)		Max Stress
(MPa)		Max Elongation
(mm)		Max Time
(sec)
306.2722.123.5352.5
314.1740.523.8279.5
310.4731.923.4251.1
309.4729.524.1244.0
307.8725.824.6241.5
307.3724.524.3233.5
Average309.2729.123.9267.0
Std Deviation2.66.00.440.8
Table 3. Compressive tests of masonry block bricks.
Table 3. Compressive tests of masonry block bricks.
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Loading Area
(mm2)		Failure Load
(kN)		Compressive Stress
(MPa)		Loading Area
(mm2)		Failure Load
(kN)		Compressive Stress
(MPa)		Loading Area
(mm2)		Failure Load
(kN)		Compressive Stress
(MPa)
36,40080.22.224,05015.90.751,800132.22.6
36,400148.94.124,05025.91.151,800164.53.2
36,400111.72.424,05036.31.551,800114.82.2
36,400120.83.324,05024.91.051,800237.64.6
36,400122.13.324,05031.91.351,8001052.0
36,400153.24.124,05047.21.951,800240.44.6
36,40095.42.624,05049.02.051,800147.12.8
36,400154.44.224,05012.80.551,800237.34.6
36,400133.63.724,05030.31.351,800147.52.8
Average124.53.3Average30.51.3Average169.63.3
StdDeviation24.50.7StdDeviation11.70.5StdDeviation51.41.0
Table 4. Compressive strength of mortar samples.
Table 4. Compressive strength of mortar samples.
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I. SeriesII. SeriesIII. Series
Loading
Area(mm2)		Failure Load(kN)Compressive Stress
(MPa)		Loading
Area
(mm2)		Failure Load
(kN)		Compressive Stress
(MPa)		Loading
Area
(mm2)		Failure Load
(kN)		Compressive Stress
(MPa)
785470.79.00785484.411.00785480.711.42
785490.711.36785479.410.11785491.911.70
785489.211.55785478.610.01785490.210.21
785493.611.92785476.39.70785493.111.85
785497.812.45785474.29.45785478.610.01
785487.911.19785474.59.49785496.112.24
785484.510.76785485.010.847854100.912.85
785483.210.59785472.59.23785480.910.30
785470.08.91785480.610.267854100.312.77
Average85.210.81Average78.410.01Average90.311.48
Std. Deviation8.91.13Std. Deviation4.20.58Std. Deviation7.981.02
Table 5. Compressive test results of glass fiber reinforced and non-reinforced shotcrete.
Table 5. Compressive test results of glass fiber reinforced and non-reinforced shotcrete.
Buildings 15 02781 i009
ShotcreteGlass Fiber-Reinforced Shotcrete (GFRC)
Loading
Area
(mm2)		Fracture
Load
(kN)		Compressive Stress
(MPa)		Elastic
Modulus (GPa)		Loading
Area
(mm2)		Fracture
Load
(kN)		Compressive Stress
(MPa)		Elastic
Modulus (GPa)
384873.155.6914.31384820753.78918.89
384872.655.3814.533848197.951.42617.99
384871.354.5913.983848203.252.79618.59
384872.355.2114.753848201.552.45718.35
Average72.355.2114.39Average202.452.61718.45
Std. Deviation0.760.46 Std. Deviation14.30.95
Table 6. Test results of the specimens used in the study.
Table 6. Test results of the specimens used in the study.
Test SpecimensLoad Carrying
Capacity
(kN)		Stiffness
(kN/mm)		Ductility RatioEnergy Absorption
Capacity
(kJ)
MW623.13.810
MWS8054.514
MWSG10565.424
MWS 2-2001567.64.637
MWS 3-4001267.45.429
MWS 3-2001679.46.439
MWS 3-1501759.96.740
MWS 3-10023010.48.252
MWSG 2-20018064.742
MWSG 1.5-1501686.16.338
MWSG 2-1501866.46.643
MWSG 3-1502057.26.947
MWSG 3-4001585.57.230
Table 7. Comparison between experimental and FEM results.
Table 7. Comparison between experimental and FEM results.
MWMWSMWSGMWS 2-200MWS 3-400MWS 3-200MWS 3-150MWS 3-100MWSG 1.5-150MWSG 2-150MWSG 2-200MWSG 3-150MWSG 3-400
Pmax-EXP (kN)98.2101197275209212219272212239255264179
Pmax-FEM (kN)98110191268218209214267217238241253173
Convergence-Dif %0.28.93.142.64.31.032.31.82.30.45.84.33.4
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MDPI and ACS Style

Yaman, Z.; Cumhur, A.; Ağcakoca, E.; Özyurt, M.Z.; Maraşlı, M.; Sadid, M.S.; Akrami, A.; Rasuly, A. Experimental Study and FEM Analysis on the Strengthening of Masonry Brick Walls Using Expanded Steel Plates and Shotcrete with and Without Glass Fiber Reinforcement. Buildings 2025, 15, 2781. https://doi.org/10.3390/buildings15152781

AMA Style

Yaman Z, Cumhur A, Ağcakoca E, Özyurt MZ, Maraşlı M, Sadid MS, Akrami A, Rasuly A. Experimental Study and FEM Analysis on the Strengthening of Masonry Brick Walls Using Expanded Steel Plates and Shotcrete with and Without Glass Fiber Reinforcement. Buildings. 2025; 15(15):2781. https://doi.org/10.3390/buildings15152781

Chicago/Turabian Style

Yaman, Zeynep, Alper Cumhur, Elif Ağcakoca, Muhammet Zeki Özyurt, Muhammed Maraşlı, Mohammad Saber Sadid, Abdulsalam Akrami, and Azizullah Rasuly. 2025. "Experimental Study and FEM Analysis on the Strengthening of Masonry Brick Walls Using Expanded Steel Plates and Shotcrete with and Without Glass Fiber Reinforcement" Buildings 15, no. 15: 2781. https://doi.org/10.3390/buildings15152781

APA Style

Yaman, Z., Cumhur, A., Ağcakoca, E., Özyurt, M. Z., Maraşlı, M., Sadid, M. S., Akrami, A., & Rasuly, A. (2025). Experimental Study and FEM Analysis on the Strengthening of Masonry Brick Walls Using Expanded Steel Plates and Shotcrete with and Without Glass Fiber Reinforcement. Buildings, 15(15), 2781. https://doi.org/10.3390/buildings15152781

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