Lateral Strength Evaluation of Ferrocement Strengthened Masonry Inlled RC frame Based on Experimentally Observed Failure Mechanisms

In developing countries, lateral strengthening of seismically vulnerable masonry infilled RC buildings is one of the major concern. In this context, ferrocement can be used as a low cost and less labor-intensive strengthening scheme for those buildings. This study aims to experimentally identify major failure mechanisms, and to develop a lateral strength evaluation procedure of ferrocement strengthened masonry infilled RC frame. Subsequently, ductility of all of the identified major failure mechanisms is compared. Mainly four major failure mechanisms (i.e. overall flexural, column punching-joint sliding, diagonal compression, and diagonal cracking-sliding) are identified from the current experimental work and past experimental studies. The strength evaluation procedure, based on the identified failure mechanisms, is proposed and verified with an average calculated to experimental lateral strength ratio of 0.8. Among the identified failure mechanisms, overall flexural, and diagonal cracking-sliding mechanisms showed relatively ductile behavior when compared to the ductility of column punching-joint sliding, and diagonal compression failure mechanism.


Introduction
In developing countries, seismic strengthening of existing buildings is one of the most important concerns for structural engineers because there are many seismically vulnerable buildings. Lateral capacity improvement of these existing vulnerable buildings could be conducted by insertion of additional structural elements e.g. shear wall, steel bracing, etc. or by increasing the lateral capacity of existing elements i.e. jacketing of RC columns, beams, walls, etc. Structural element insertion would be difficult to apply in several thousands of existing RC buildings of developing countries due to limitation of expertise and capital. On the other hand, jacketing technology might be a feasible approach for developing countries, if it could be conducted with less technical efforts and capital. In many developing countries, the common structural system of buildings is RC moment resisting frame with non-structural masonry infill walls which might be a probable candidate for strengthening. Some reinforcing material is required to convert the existing infill masonry to structural components with an intension to attain a reliable performance during an earthquake. The approach of converting infill masonry to structural components might serve two purposes. First, it would be a strengthening method and second, it might prevent the un-expected low strength of masonry infilled RC frame caused by short column effect by sliding of infill masonry. Strengthening of existing infill masonry can be conducted by Ferrocement (FC) lamination, Textile mortar reinforcement (TRM), Fiber reinforced mortar (FRM), etc. Among these, ferrocement lamination is a low cost and less labor intensive strengthening method for developing countries. As a part of SATREPS-TSUIB project in Bangladesh (https://www.satreps-tsuib.net/), which is sponsored by JICA (Japan International Cooperation Agency), authors are trying to develop an effective evaluation procedure of ferrocement strengthened masonry infilled RC frame.
Ferrocement strengthening of masonry refers to the application of an initial mortar layer on both surfaces of the masonry wall which is followed by the placement of steel mesh reinforcement and a second mortar layer, as shown in Fig. 1. Anchorages are also being used to attach wire-mesh to masonry and RC frame. Hereafter, the term "Ferrocement" will be stated as "FC" throughout the article.
Though FC has been studied for decades as a construction material, there is no design specification e.g. amount of mesh reinforcement, mortar thickness, etc. for use as a shear strengthening material on unreinforced infilled masonry. The paucity of comprehensive studies on FC strengthening has also been acknowledged in ACI 549 (1997). Though, several researchers (Kaya et al. 2018, Seki et al. 2018, Demirel et al. 2015, Altin et al. 2010, Zarnic and Tomazevic 1985 studied the effect of FC on the lateral behavior of masonry infilled RC frame, however, the main intention was to improve seismic capacity rather than to identify and evaluate experimentally observed failure mechanisms. Therefore, this study initially aims to recognize major failure mechanisms of masonry infilled RC frame with ferrocement strengthening based on a new experimental campaign (with 2 half-scaled specimens) as well as from past experimental studies. Then, endeavor is taken to propose and validate lateral strength evaluation of the FC strengthened masonry infilled RC frame based on the recognized major failure mechanisms. Finally, experimental ductility level of the major failure mechanisms is compared.

Identified load transfer mechanisms in past experimental studies
In general, seismic performance evaluation of a structural component needs a proper understanding of load transfer mechanism i.e. failure mechanism of that structural component. Therefore, an attempt is taken here to understand load transfer mechanisms of masonry infilled RC frame after ferrocement strengthening. Since ferrocement is to be used for strengthening of masonry infilled RC frame, it is also necessary to understand the failure mechanisms of un-strengthened masonry infilled RC frame prior. It will help to anticipate the most probable failure mechanisms of masonry infilled RC frame after ferrocement strengthening.
Generally, the failure mechanism of masonry infilled RC frame depends on relative stiffness as well as the strength of infill material compared to the surrounding RC frame (ASCE/SEI 41-06 2001).
At lower lateral displacement, masonry infilled RC frame works similar to a monolithic infilled frame.
Then, with the increasing lateral displacement, a separation occurs between frame and infill masonry, subsequently, the RC frame acts like a braced RC frame. Before separation, the infilled RC frame can also fail by overturning or rocking like a flexural wall when the infill masonry is very strong and stiff.
Overall flexural failure (i.e. rocking) can happen when the RC column is lightly reinforced (FEMA 306 1998) and this failure has been reported in a study by Adnan et al. (2019). After separation of masonry and the surrounding frame, if the infilled RC frame can escape overall flexural failure, the lateral load is carried by both components therefore both RC frame and masonry experience damages. In general infill masonry exhibits diagonal compression crushing, joint sliding, and/ or diagonal cracking, whereas RC columns fail by the formation of plastic hinges or shear failure of RC columns. In some of the cases, especially when sliding occurs at the mid-height, a short column is also evident (Zahura et al. 2020 andAlwashali et al. 2018 Zahura et al. (2020) and Alwashali et al. (2018), therefore the combination of sliding and diagonal cracking is considered herein as one of the failure mechanism. However, when the sliding plane is at the interface of beam soffit and upper infill masonry course, it might also be possible to get an extremely short column failure of the tension column that might fail by horizontal slip shear failure, as shown in Fig. 2 (Kaya et al. 2018, Seki et al. 2018, Demirel et al. 2015, Altin et al. 2010, Zarnic and Tomazevic 1985 in order to understand the major load transfer mechanisms i.e. failure mechanisms of ferrocement strengthened masonry infilled RC frame. In those studies (Kaya et al. 2018, Seki et al. 2018, Demirel et al. 2015, Altin et al. 2010, Zarnic and Tomazevic 1985, both solid clay bricks and hollow bricks, having masonry compressive strength of 5 ~ 15MPa, were used to build infill masonry. The masonry infills were strengthened using ferrocement having a mortar strength of 3 ~ 30MPa and wire mesh ratio of 0.05 ~ 0.35%. Kaya et al. (2018) conducted an experimental investigation having seven FC strengthened infill hollow brick masonry specimens. The main variables were the embedment length and spacing of the dowel bar, where dowel bars have been used to transfer shear from the RC frame to the FC layer. It is to be noted that FC has been applied on one side of the infill panel only. The strengthened surface experienced a distributed shear cracks pattern and the failure mechanism has been recognized as diagonal compression mechanism. The experimental investigation indicated that the dowel length and spacing did not affect the maximum resistance which is on average 1.5 times when compared to the unstrengthened masonry infilled RC frame.

Fig. 2
Extremely short column i.e. punching failure of tension column (Crisafulli 1997)  rather than the connection of mortar to the RC frame. It is worthy to note that sometimes delamination was also happened (Kaya et al. 2018), however delamination is not considered as a major failure mechanism i.e. load transfer mechanism.
In past studies, mainly two major load transfer mechanisms i.e. diagonal compression (Kaya et al. 2018 andDemirel et al. 2015), and diagonal cracking-sliding (Seki et al. 2018, Altin et al. 2010, Zarnic and Tomazevic 1985 were evident where both FC strengthened infill panel and surrounding RC frame damaged simultaneously. This could be attributed to the low strength masonry as well as FC mortar. If infill masonry and strengthening mortar had higher strength it could have more impact on the surrounding frame with concentrated damage of the surrounding RC frame only like an overall flexural failure, and column punching-joint sliding failure that have been found for un-strengthened masonry infilled RC frame. Since, overall flexural failure, and column punching-joint sliding failure are not evident in past studies, therefore it is necessary to investigate ferrocement strengthening on strong masonry infill i.e. high compressive strength. This is the main motivation for the experimental investigation on ferrocement strengthening of strong masonry infilled RC frame which is discussed in the following section. At the same time, past experimental studies focused on the mortar strength and connection of FC with the RC frame as governing parameters of FC strengthened masonry infilled RC frame. However, the amount of wire mesh has not been studied yet, therefore wire mesh ratio is set as the variable for test specimens.

Specimen design concept
Two half-scaled ferrocement strengthened masonry infilled RC frames, as shown in Fig. 3(a)-(d), are designed to be in line with the objective of this study. The design aspect of each component is discussed below:

RC frame
As mentioned earlier, this study is focused to be applied in Bangladesh therefore a cross-section of 200 x 200mm of the RC column has been adopted as a half scaled dimension to represent field practice in Bangladesh. The long and shear reinforcements of the RC columns have been designed as a flexural column. The top beam was intentionally kept very stiff to avoid failure on the beam.

Infill masonry
Strong masonry infill (masonry compressive strength > 25MPa) was built inside the RC frame for

Ferrocement
Several types (i.e. square, hexagonal, diamond shape) of mesh reinforcements are available which can be used for FC strengthening of masonry. In this study, steel wire mesh having square spacing is adopted for infill masonry strengthening because square wire mesh is recommended by ACI 549R (1997) to use where the biaxial stress field is prominent e.g. masonry infill. In this study, wire mesh content is selected based on past literature survey, since there is no design guideline for ferrocement strengthening of masonry infilled RC frame. In past studies (Kaya et al. 2018, Seki et al. 2018, Demirel et al. 2015, Altin et al. 2010, Zarnic and Tomazevic 1985, horizontal mesh reinforcement was kept within a range of 0.05~0.35% of the horizontal masonry area. Since the objective of this experimental program is to investigate the impact of infill panel on the surrounding RC frame, therefore wire mesh ratio for the current study has been set to 0.16% (relatively low ratio) for specimen IM-FC-1 and 0.56% (relatively high ratio) for specimen IM-FC-2. For specimen IM-FC-2, wire mesh ratio has been set almost 3.5 times than specimen IM-FC-1, to investigate the effect of higher wire mesh ratio on the failure mechanism of the surrounding RC frame.

Connection with RC frame and masonry
In general, the connection of FC layer with the RC frame can be established by dowel bars. However, Kaya et al. (2018) concluded that dowel spacing and dowel embedment length might not have an influence on the lateral strength of FC strengthened masonry infilled RC frame. In addition, Altin et al. (2010) found that the dowel bar causes premature bearing failure of the FC mortar layer without much improvement in lateral capacity in case of low strength mortar which is commonly used in developing countries. One additional point to be noted that very stiff connection, i.e. dowel bar, between mortar and RC frame might cause easy splitting of the applied FC mortar layer in case of low strength mortar since no spiral reinforcement is provided. Therefore, the dowel connection is not considered herein.
The insertion of mesh reinforcement in the RC frame is effective than dowel connection between the FC layer and RC frame as reported by Altin et al. (2010). However, for practical strengthening purpose, it is complicated to insert the mesh reinforcement into the RC frame. Therefore, in this study, a new idea has been conceived to secure a connection of the RC frame to wire mesh. The main intension was to make a good attachment of wire mesh to the RC frame, so that wire mesh can actively participate in load transferring. To execute that steel plate and bolt connection has been considered for this study as shown in Fig. 3(c)-(d). The spacing of the bolts has been kept smaller near the RC beam-column joint considering the commencement of separation at that region which might occur without any connection.
12 bolts (8mm in diameter), as shown in Fig. 3(b), have been used along with steel plate with mixed spacing, 100 and 200mm between bolts.
In ferrocement strengthening, wire mesh should also be connected to masonry panel in order to hold it during construction as well as to avoid delamination under lateral loading. Alcocer and Flores (2001) investigated the quantity of optimum connection for ferrocement strengthening of confined masonry and recommended about nine connectors per square meter that have been adopted in this study.

Test specimen details
Two half-scaled masonry infilled RC frames have been constructed and infill masonry has been strengthened with ferrocement. The overall geometry of the RC frame is shown in Fig. 3(a). The details of both specimens are shown in Table 2. The construction procedure of specimens is as follows: First, the RC frame has been constructed and then masonry panel has been built inside the frame, with solid bricks of 210x100x60 mm, in running bond manner. After seven days of masonry construction, 10mm thick mortar has been mounted on both faces of the masonry wall. This is followed by the attachment of square wire mesh to the RC frame and masonry wall. The wire mesh has been connected to the surrounding RC frame with bolts (inserted into the pre-installed thread) and steel plate as shown in Fig. 3(d). In addition, the wire mesh has been connected with masonry infill by 32mm long nails to hold the wire mesh in place during the application of second layer mortar. The nails have been placed in drilled holes with glue at a horizontal and vertical center to center distance of 250mm and 500mm, respectively. After seven days, the second layer of mortar, having 15mm thickness, has been applied on the wire mesh. Therefore, FC mortar thickness is about 25mm on each surface of infill masonry.

Material properties
The material tests of concrete, reinforcing steel have been conducted for each specimen as per Japanese standard (2010). The wire mesh has been tested as per ACI 549R (1997). The masonry compressive strength has been tested according to ASTM (2011). The mechanical properties of concrete, reinforcing steel, masonry, mortar, and wire mesh are shown in Table 3.

Test setup and instrumentation
Several strain gauges were attached to the long and tie reinforcement of RC column to measure the strain during loading. The locations of the strain gauges are shown in Fig. 4   lateral cyclic loading is shown in Fig. 6(f) and Fig. 8. It is to be noted that no other crack, except at the bottom horizontal crack, was observed on the FC strengthened masonry infill. In addition, there was no crack at the top construction joint.

Response under cyclic lateral load Specimen IM-FC-1
The hysteresis loops of IM-FC-1 is shown in Fig. 9(a). The response was essentially linear up to the formation of the first crack on the tension column at 0.05% story drift. After cracking, the hysteresis loops began to wide, specifically after 0.4% story drift when an inclined crack appeared on the ferrocement laminated masonry. The peak resistances were +538kN and -530kN at the story drift of +0.39% and -0.38, respectively. After peak resistance, the stiffness of the hysteresis loop gradually increases with load enhancement, exhibiting a "pinching" in the hysteresis curve. This could be attributed to the sliding at the top joint as discussed before. At around 0.6% drift, a sudden drop in lateral resistance occurred following the sliding at the top construction joint. As the sliding continues, the residual resistance of the specimen became almost constant which is marked by the horizontal postpeak branch of the envelope curve as shown in Fig. 9(a). The story drift capacities at residual strength are 0.55 and 0.56 in positive and negative loading directions, respectively. The residual capacity is considered at the 80% of peak resistance on the post-peak branch. The hysteresis loops of IM-FC-2 is presented in Fig. 9(b). The behaviour was essentially elastic until the formation of the initial crack on the tension column at 0.05% story drift. After cracking, the relatively stiffer hysteresis loops continued until the specimen IM-FC-2 reached to the peak resistance of +593kN and -583kN at the story drift of +0.

Recognition of failure mechanism
The failure mechanism i.e. load transfer mechanism at peak resistance is necessary to be identified for lateral strength evaluation that is to be discussed in the next section. The aforementioned crack propagation and hysteresis behaviour, indicate that specimen IM-FC-1 undergone both flexural and shear damages whereas IM-FC-2 mostly undergone through flexural behaviour throughout the courses of lateral story drift. However, the following analysis were conducted to recognize the load transfer mechanism at the peak resistance.

Shear and flexural component of story deformation
The flexural and shear components of total lateral deformation were measured to confirm the failure modes of the specimens. The flexural component (∆flex) of lateral story deformation was measured using the LVDTs attached over the height of RC columns as shown in Fig. 4 Flexural deformation, Δflex =∫ ℎ (1) Shear component of lateral story deformation, Δsh = Δstory -Δflex The obtained flexural and shear deformation in relation to the story drifts are shown in Fig. 10(a)-(b). In specimen IM-FC-1, the flexural contribution is relatively more at lower story drifts as shown in Fig. 10(a). At higher story drifts, the tension column experienced punching shear failure following sliding at the top joint which led to an increase in shear deformation. Another strengthened RC frame, namely IM-FC-2, experienced flexure domination throughout the course of the lateral drift as shown in Fig.10(b). In other words, IM-FC-1 behaved as a flexural wall at the drift lower than 0.4% and then failed in punching shear of the tension column, however the specimen IM-FC-2 behaved like a flexural wall for all story drifts.

Rotation at the top of tension column
In specimen IM-FC-1, at higher story drift level the tension column was failed in punching shear where the long reinforcement bent at the level of tie reinforcement as shown in the close view of Fig. 7.
Therefore, the rotation of four column sections, where LVDTs were attached, on the upper half of the tension column are investigated as shown in Fig. 11(a). It is obvious that until 0.2% story drift there was almost no rotation at the column sections on the upper half of the tension column. Then, rotation at the top column section, at level 1380mm from the base, increased gradually leading to the reinforcement bent i.e. punching shear failure at around 1.5% story drift. In specimen IM-FC-2, there was no visible sliding at the top construction joint, as discussed earlier, therefore no trace of punching shear failure i.e. long reinforcement bent was found. However, the rotation at the aforementioned four column sections on the upper half of the tension column is investigated as shown in Fig. 11(b) to confirm the fact. It is obvious that there was almost no rotation at the top column section at level 1380mm from the base, hence no punching shear damage of the tension column. From the above two analyses, it is clear that the specimen IM-FC-1 started to rock at lower story drifts levels, until 0.4% story drift, like a structural wall with boundary columns. At higher story drifts, the rocking behavior was ceased due to the sliding at the top construction joint which ultimately leads to punching shear failure of the tension column. Therefore, the load transfer mechanism is idealized as an overall flexural failure at peak resistance. On the other hand, specimen IM-FC-2 started to rock from the beginning and continued rocking behavior until failed with rupture of long reinforcement of the tension column. Therefore, the failure is idealized as fully developed overall flexural failure.

Recognized major failure mechanisms
As discussed earlier, the failure mechanism of the infilled RC frame depends on relative stiffness as well as the strength of infill material compared to the surrounding RC frame. When relatively weak masonry infill is strengthened with ferrocement, failure initiates on the infill panel by diagonal compression crushing (Kaya et al. 2018 andDemirel et al. 2015) or diagonal cracking and/or sliding (Seki et al. 2018, Altin et al. 2010, Zarnic and Tomazevic 1985 as recognized from past experimental studies. However, when strong masonry is strengthened with ferrocement, the RC frame could be the weakest part and consequently can fail by overall flexural or column punching leaving the infill panel Based on the current experimental observation and previous studies four distinct major failure mechanisms, as shown in Fig. 12, are recognized.

Lateral strength evaluation of each failure mechanism
Overall flexural (Failure I) Overall flexural failure is likely to occur when the infill panel (masonry and FC layer) is very strong and stiff. This phenomenon can be attributed to the fact that when the infill panel is very strong and stiff, it cannot deform in a shear mode therefore the structural system undergoes a rocking mode like a rigid body. This rocking behavior generates tensile yielding of the tension column longitudinal reinforcements followed by failure due to rupture of longitudinal reinforcement of the tension column (left side close view in Fig. 13), and concrete crushing at the bottom of the compression column (right side close view in Fig. 13).
The idealized load transfer mechanism by overall flexural yielding is shown in Fig. 13   Generally, punching shear failure occurs at the top end of the tension column, as explained prior in Fig. 14, within a very short distance from the face of the top beam. Therefore, this type of shear failure is known as an extremely short column (JBDPA 2001) where the shear span ratio is less than unity. Since the shear span ratio is very low, hence the shear failure is considered alike to slippage (Yamamoto 1990) where shear resistance comes more likely from shear friction rather than shear reinforcement. The punching shear capacity (psQc) of the column is determined based on the procedure suggested by JBDPA (2001). In JBDPA (2001), punching shear capacity is evaluated by Eq. (7). In evaluation, the basic shear strength (τo) is computed, using Eq. (8) -Eq. (10), based on shear friction (Yamamoto 1990) i.e. taking into account the influence of long reinforcement and axial stress level on the shear plane. However, to include the influence of shear span ratio (a/d), an influence factor (kmin) has also been considered in Eq. (7), as suggested by Yamamoto 1990 as per Eq. (11). The shear span ratio (a/d) is considered equal to 1/3 for all the specimens as suggested by JBDPA (2001)  The source of joint shear capacity (jsQw) could be the masonry joint mortar, mortar of the FC layer, and embedded wire meshes in the FC layer. From the lateral behavior of specimen IM-FC-1, it is clear that initially shear strength is greater than flexural capacity (i.e. shear failure after flexural yielding) and the lateral resistance degraded 25% after occurring slippage at the top construction joint. This indicates that initially, the bond between FC laminated masonry and soffit was working. Then, after slippage, wire meshes were working as a dowel to provide residual capacity. Therefore, at the initial stage, before any slippage at the interface of infill top and soffit, shear capacity can be considered as shear strength (cohesion) of mortar at the interface. In initial bond capacity, wire mesh might have a contribution in addition to mortar cohesion however, as a conservative approach wire mesh contribution is ignored.
After the occurrence of slippage wire mesh will be subjected to shearing force hence can be considered as the source of residual shear capacity at the interface. The joint shear capacity can be evaluated from Eq. (13).
where, jsQw = shear capacity at joint; τmas, τmor,FC = shear strength (cohesion) of mortar in masonry joint and ferrocement; lw = length of infill; tmas , tFC = thickness of masonry wall and FC layer; and ns = number of FC surface. It is to be noted that cohesion capacity of mortar, for both masonry and FC layer, has been considered as 0.17√fmor (where fmor =compressive strength of mortar), which has been recommended by Naaman (2000), and Mander and Nair (1994) as shear strength of FC.

Diagonal compression (Failure III)
In diagonal compression failure, the infilled part (masonry and FC layer) is considered to behave similarly to a diagonal strut, as shown in Fig. 15 where, Ws = strut width; ac = contact length between RC column and infill panel; and θ = inclination of loaded diagonal with horizontal. The contact length is calculated by considering that the RC column is resting on an infill panel, as shown in Fig.16(a), similar to a beam on an elastic foundation (Hetenyi 1946). Therefore, lateral deflection (y) and curvature (φ) of the RC column can be evaluated considering it is analogous to a beam on an elastic foundation, as shown in Fig. 17(b) and 17(c). Relative rigidity (λ) of infill panel with respect to RC column is defined as Eq. (16).
where, λ = relative rigidity factor for FC strengthened masonry, θ = inclination of loaded diagonal with horizontal; Ec, Emas, EFC = Young's modulus of concrete, masonry, FC mortar; tmas, tFC = thickness of masonry, FC mortar layer; ns = number of surface retrofitted with FC; Ic = moment of inertia of RC column; and dm = diagonal length of infill panel.
It is evident from the column deflection shape, as shown in Fig. 16(b), that the lower portion of the RC column exhibits flexure deflection whereas the deflection mode of the upper part is changed from a flexural shape due to the presence of infill masonry, which actually causes the separation between masonry and RC frame. Based on the deflection shape, it is considered that the infill panel of = 4 (17)

Diagonal cracking and sliding (Failure IV)
In case of diagonal cracking and sliding failure, at lower story drifts level, a diagonal tensile crack appeared on FC strengthened infill panel, as shown in Fig. 17(a). However, at post-peak stage the strengthened masonry can have a sliding behavior as shown in Fig. 17(b). Therefore, this failure mechanism is defining as diagonal cracking and sliding failure. Since peak resistance occurs at the diagonal cracking stage, therefore the idealization of the load transfer mechanism at the peak is considered as shown in Fig. 18. The FC strengthened masonry cracked on the loaded diagonal direction whereas the frame behaves like a bare frame. At diagonal cracking, lateral strength can be evaluated, as a summation of the strength of RC columns, masonry, and wire meshes of ferrocement, using by Eq. (18).
where, fQc = lateral capacity of RC column (can be evaluates from Eq. (11)  The contribution of infill masonry due to diagonal cracking has been evaluated by the second term of Eq. (18), in reference to Fig. 18(b). Masonry cracking strength (fmas,cr) is considered as 5% of masonry compressive strength (Sen 2020), which has been proposed based on diagonal wallet test results. The contribution of ferrocement has been considered as the shear capacity provided by the horizontal wire meshes as the third term of Eq. (18). The contribution of FC mortar layer in cracking capacity is not considered assuming the fact that the mortar and wire mesh will not work together because FC layer mortar would crack before wire meshes come into tension. Conceptually, the contribution of horizontal wire meshes is assumed like reinforcements in a RC wall. However, Alcocer et al. (2003)  might not reach to yield condition. Assuming the similar behavior for the ferrocement strengthened masonry, an empirical reduction factor (α) has been imposed in Eq. (18) to accommodate the less effectiveness of mesh reinforcement compared to contribution in the RC shear wall. The less effectiveness might happen due to the less development length provided for reinforcements in masonry. Anderson and Priestley (1992) observed that the strength of shear reinforcement in a reinforced masonry wall is approximately half than that of in RC shear wall. Several building codes e.g. MSJC (2011); CSA-S304.1 (2004) andNZS 4230 (2004) have also taken account less effectiveness of horizontal reinforcement for internally reinforced masonry. The range of the reduction factor varies in between 0.5 ~ 0.8. In this study, the empirical reduction factor (α) has been assumed as 0.7 for ferrocement lamination as suggested in Sen et al. (2020).

Proposal and validation on lateral strength evaluation
Since, masonry infilled RC frame with ferrocement strengthening can undergo several load transfer mechanisms, as discussed in the earlier section, therefore it is important to recognize the expected load transfer mechanism which likely to occur under lateral loading at seismic events. In general, the expected load transfer mechanism should be the one with the lowest lateral resistance. Therefore, the expected lateral strength (Qcal) will be the minimum of the capacities of all load transfer mechanisms as shown in Eq. (19).
All the calculated lateral capacities and experimental capacities are presented in Table 4. In the case of overall flexural (specimen IM-FC-2), diagonal compression (specimen Sp-5 (Kaya et al. 2018) and SMF (Demirel et al. 2015)), and diagonal cracking-sliding (specimen S5-FMFC (Seki et al. 2018), Sp-4 (Altin et al. 2010), and M3 (Zarnic and Tomazevic 1985)) load transfer mechanism, the minimum calculated capacity can predict the failure mechanism. In case of mixed failure (specimen IM-FC-1) and column punching-top joint sliding failure (specimen AR ) and IM-FC (Zahura et al. 2020)) the predicted failure mechanism deviates from the actual failure mechanism. Therefore, it can be concluded that in most cases the proposed evaluation method can predict the failure mechanism.
In addition, the calculated lateral strength (Qcal) and experimental lateral strength (Qexp) for all the investigated specimens are plotted in Fig. 19. In almost all the cases the prediction is on the safe side. The average calculated to experimental capacity ratio (Qcal / Qexp) is 0.8 with a coefficient of variation 0.2.

Comparison of ductility level for all failure mechanisms
Ductility of a structure indicates the capability of that structure to undergo large inelastic deformation without a substantial reduction in lateral strength. There are several parameters, e.g. absorbed energy and lateral displacement, to define ductility. In this study, displacement ductility factor (μ) ductility is considered as the ratio of ultimate and yield lateral story drift as shown in Eq. (20).
Under lateral load, structure does not always show a well-defined yield point, therefore there are several ways to consider yield point for a structure under lateral load. Park (1988) summarized four commonly used ideas to define yield deformation of a structure in laboratory test. The idea of equivalent bilinear elasto-plastic yield point with reduced stiffness, as shown in Fig. 20, is recommended by Park (1988) since it considers deterioration of stiffness because of cracking and this stiffness deterioration is usual for real structure. Therefore, this idea is adopted in this study to define the experimental yield point (Ry) of the investigated test specimens at 75% of ultimate strength (i.e. 0.75Qmax) as presented in Fig. 20.
Most of the structures have some deformation capability beyond the peak resistance. Therefore, it is appropriate to consider the ultimate deformation capacity at the post-peak stage where load-carrying capacity reduces substantially (Park 1988). In this study, the ultimate deformation capacity (Ru) of the test specimen is considered at 20% lateral strength drop after peak resistance as shown in Fig. 20. All of the computed yield drifts and ultimate drifts are summarized in Table 5  The experimental ductility factor (μexp) of all the investigated specimens are shown on normalized lateral load-displacement curve in Fig. 21. Overall flexural mechanism (Failure-I) exhibited a relatively high ductility (μexp =14) than other failure mechanisms. On the other hand, column punching and joint sliding failure (Failure II) showed relatively low ductility (μexp = 3.4 ~ 4), when compared to other failure mechanisms. Meanwhile, the diagonal compression mechanism (Failure-III) showed relatively higher ductility (μexp = 4.6 ~ 6.7) than the column punching-joint sliding mechanism (Failure-II), however less than diagonal cracking and flexural failure. Diagonal cracking and sliding mechanism (Failure-IV) showed ductility within a wide range of 3.4 ~10. The comparison of experimental ductility factors is also shown in Fig. 21, from which it is evident that overall flexural and diagonal crackingsliding mechanism can be treated as ductile behavior. Whereas, column punching-joint sliding, and diagonal compression can be considered as a relatively brittle failure. o Overall flexural and diagonal cracking-sliding mechanisms exhibit relatively ductile behavior.
Whereas, column punching-joint sliding, and diagonal compression mechanism can show a relatively brittle failure.
These conclusions are based on very limited number of test specimens; therefore, more studies are required before application of this evaluation procedure in real field. Figure 1 Schematic diagram of FC strengthened masonry in ll   Major failure mechanisms of FC strengthened masonry in lled RC frame Figure 13 Load transfer mechanism of overall exural failure (Failure I) Figure 14 Idealized load transfer mechanism of column punching and joint failure (Failure II)

Figure 15
Idealized load transfer mechanism of diagonal compression failure (Failure III) Figure 16 Surrounding RC column (a) un-deformed shape, (b) deformed shape and (c) curvature distribution   Experimental ductility factor of all investigated test specimens