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The monolithic structural action of various types of walls of flanged cross section and walls with engaged stiffeners or returns (nonrectangular sections) is critically dependent on the shear capacity of the interface between the components making up the section. An assessment of the shear capacity of the interface may also be required as part of t...
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Context 1
... and experimental studies of brick masonry wall strength generally have concentrated on the consideration of rectangular cross sections. However, many walls in practical situa- tions are stiffened by piers, returns, or other flanged sections in order to increase their lateral resistance ͑ Fig. 1 ͒ . To achieve effective composite behavior for these types of nonrectangular walls, monolithic structural connection is required across the vertical interface between the masonry components. The supporting flanges are usually linked to the web by masonry bond in the form of header courses or by metal shear connectors embedded in the bed joints and extending across the shear plane with the vertical joint at the interface filled with mortar. This latter technique is useful where there is a difference in the bonding pattern on each side of a wall or where different materials are used side by side. Connectors are also useful when a damp-proof membrane is in- corporated at the vertical joint between the ribs and the outer leaf of a diaphragm wall to prevent the passage of moisture in highly absorbent units. In design, insufficient vertical shear capacity along the interface may be the result of significant wall section changes down the building or the effects of differential vertical loading on the flanges and web of the section ͑ Sinha and Hendry 1981; Correa and Ramalho 2003 ͒ . These effects, together with uneven settlements of the structure, or more often wind or seismic actions ͑ Drysdale et al. 1994, 2008 ͒ are just some of the examples where the shear capacity of the interface in nonrectangular sections could be the critical parameter in design. This interface capacity may be relevant to elastic design ͑ depending on code requirements ͒ , and it is certainly relevant to ultimate strength design where inelastic effects must be considered. This problem will also be of interest to designers in the case of extreme loading, where masonry structures may be required to have enhanced structural integrity as part of a comprehensive design strategy against pro- gressive collapse due to accident, misuse, sabotage, or other causes. For laterally loaded nonrectangular cross sections, for the calculation of the section modulus of the shear wall ͑ Fig. 1 ͒ , the flange portion of a shear wall to be considered is usually restricted to within certain limits of l / h ratio to allow for shear lag effects ͑ l —length of the wall and l , h —height of l / h the Յ 1.5 wall ͒ . Flange requirements for C-, L-, l e = and Ά 0.75 Z-shaped h + 0.5 l , sections 1.5 Յ l are / h Յ usually 3.5 · limited up ͑ 1 to ͒ 1/6 of the total height of 2.5 the h , wall or six l / h times Ն 3.5 the thickness of the intersecting These equations wall, may whichever result in is effective smaller. flange For T-section widths greater walls than the width 1.5 times of flange the height sections or three is usually quarters double of the the wall above. height Note on that either the design side of the limits web for of flanged the wall sections and will may result differ in a significant depending increase on the design in the flexural code ͓ Australian resistance Standard of the nonrectangular ͑ 2002 ͒ ; EN ͑ masonry 2005a ͒ ; Masonry element. Standards Since the new Joint MSJC Committee standard ͑ MSJC uses ͒ elastic 2008; design Canadian for uncracked Standards Association unreinforced ͑ CSA masonry ͒ 2004; ͑ URM etc. ͒ ͔ . ͑ Fig. 2 According ͒ , the to inherent some authors upper limit ͑ Orton on 1986 ͒ , in higher buildings if the flange portion is taken into ac- count, it is prudent to limit the effective width of this flange portion of the shear wall to half of that assumed for a wall in local bending—i.e., the effective width of the flange of T- and I-shaped walls would be about h / 6 and that of C-, L-, and Z-shaped walls about h / 12. Usually shear stresses along the line connecting the flange to the web should be checked if the flange portion exceeds about 40% of the length of the web. Note that according to the new MSJC ͓ Masonry Standards Joint Committee ͑ MSJC ͒ 2008 ͔ , the effective length of flange ͑ l e ͒ , which depends also on the l / h ratio of the web element is calculated as l , l / h Յ 1.5 l e = Ά 0.75 h + 0.5 l , 1.5 Յ l / h Յ 3.5 · ͑ 1 ͒ 2.5 h , l / h Ն 3.5 These equations may result in effective flange widths greater than 1.5 times the height or three quarters of the wall height on either side of the web of the wall and will result in a significant increase in the flexural resistance of the nonrectangular masonry element. Since the new MSJC standard uses elastic design for uncracked unreinforced masonry ͑ URM ͒ ͑ Fig. 2 ͒ , the inherent upper limit on flexural capacity imposed by these assumptions will usually result in noncritical levels of shear stress at the flange-web interface. This may not be the case if a cracked section analysis is performed. Only a limited amount of research has been carried out in this area. As a consequence, code design rules vary considerably from country to country and reflect the limited knowledge available. In-depth studies of this phenomenon have been carried out in relation to diaphragm wall behavior by Phipps in the United Kingdom ͑ Phipps and Montague 1986 ͒ , and this work was then extended in Australia by Phipps and Page ͑ 1995a,b ͒ . More re- cently, this problem has been evaluated also by Correa and Ramalho ͑ 2003 ͒ and Drysdale et al. ͑ 2008 ͒ . This general lack of knowledge of vertical shear capacity has resulted in variable and conservative design provisions and excessively conservative strength predictions are thus obtained when these code procedures are used in the design of load bearing structures. Two of the complicating factors in developing harmonized design provisions are the widely varying nature of wall types and construction practices and detailing in various countries. In some code provisions, the theoretical parabolic stress distribution ͑ f v = VQ / I n b ͒ is used to calculate the shear stress rather than the average stress approach ͑ f v = V / A ͒ . However, many codes use average shear stress, so direct comparison of allowable values according to different code provisions is not always valid. For the purpose of this paper European Union, Australian, Canadian, and U.S. codes are summarized and compared in Table 1. In the latest version of the new European Standard for masonry structures ͑ EN 2005a ͒ the vertical shear capacity of the connection between two adjacent walls is only very briefly dis- cussed. The vertical shear resistance of the junction of two masonry walls is required to be obtained from suitable tests ͑ which at the moment are not defined ͒ . In the absence of test results, the characteristic vertical shear resistance may be based on the initial horizontal shear resistance under zero compressive stresses ͑ f v k 0 ͒ . This contrasts with the Australian Standard for masonry structures ͑ Australian Standard 2002 ͒ which has quite specific ͑ but probably conservative ͒ provisions ͑ Table 1 ͒ . It also has some unique requirements for the shear capacity of individual connectors. These were based on a mechanistic approach and suitable for limit design methods. The U.S. design ...
Context 2
... and experimental studies of brick masonry wall strength generally have concentrated on the consideration of rectangular cross sections. However, many walls in practical situa- tions are stiffened by piers, returns, or other flanged sections in order to increase their lateral resistance ͑ Fig. 1 ͒ . To achieve effective composite behavior for these types of nonrectangular walls, monolithic structural connection is required across the vertical interface between the masonry components. The supporting flanges are usually linked to the web by masonry bond in the form of header courses or by metal shear connectors embedded in the bed joints and extending across the shear plane with the vertical joint at the interface filled with mortar. This latter technique is useful where there is a difference in the bonding pattern on each side of a wall or where different materials are used side by side. Connectors are also useful when a damp-proof membrane is in- corporated at the vertical joint between the ribs and the outer leaf of a diaphragm wall to prevent the passage of moisture in highly absorbent units. In design, insufficient vertical shear capacity along the interface may be the result of significant wall section changes down the building or the effects of differential vertical loading on the flanges and web of the section ͑ Sinha and Hendry 1981; Correa and Ramalho 2003 ͒ . These effects, together with uneven settlements of the structure, or more often wind or seismic actions ͑ Drysdale et al. 1994, 2008 ͒ are just some of the examples where the shear capacity of the interface in nonrectangular sections could be the critical parameter in design. This interface capacity may be relevant to elastic design ͑ depending on code requirements ͒ , and it is certainly relevant to ultimate strength design where inelastic effects must be considered. This problem will also be of interest to designers in the case of extreme loading, where masonry structures may be required to have enhanced structural integrity as part of a comprehensive design strategy against pro- gressive collapse due to accident, misuse, sabotage, or other causes. For laterally loaded nonrectangular cross sections, for the calculation of the section modulus of the shear wall ͑ Fig. 1 ͒ , the flange portion of a shear wall to be considered is usually restricted to within certain limits of l / h ratio to allow for shear lag effects ͑ l —length of the wall and l , h —height of l / h the Յ 1.5 wall ͒ . Flange requirements for C-, L-, l e = and Ά 0.75 Z-shaped h + 0.5 l , sections 1.5 Յ l are / h Յ usually 3.5 · limited up ͑ 1 to ͒ 1/6 of the total height of 2.5 the h , wall or six l / h times Ն 3.5 the thickness of the intersecting These equations wall, may whichever result in is effective smaller. flange For T-section widths greater walls than the width 1.5 times of flange the height sections or three is usually quarters double of the the wall above. height Note on that either the design side of the limits web for of flanged the wall sections and will may result differ in a significant depending increase on the design in the flexural code ͓ Australian resistance Standard of the nonrectangular ͑ 2002 ͒ ; EN ͑ masonry 2005a ͒ ; Masonry element. Standards Since the new Joint MSJC Committee standard ͑ MSJC uses ͒ elastic 2008; design Canadian for uncracked Standards Association unreinforced ͑ CSA masonry ͒ 2004; ͑ URM etc. ͒ ͔ . ͑ Fig. 2 According ͒ , the to inherent some authors upper limit ͑ Orton on 1986 ͒ , in higher buildings if the flange portion is taken into ac- count, it is prudent to limit the effective width of this flange portion of the shear wall to half of that assumed for a wall in local bending—i.e., the effective width of the flange of T- and I-shaped walls would be about h / 6 and that of C-, L-, and Z-shaped walls about h / 12. Usually shear stresses along the line connecting the flange to the web should be checked if the flange portion exceeds about 40% of the length of the web. Note that according to the new MSJC ͓ Masonry Standards Joint Committee ͑ MSJC ͒ 2008 ͔ , the effective length of flange ͑ l e ͒ , which depends also on the l / h ratio of the web element is calculated as l , l / h Յ 1.5 l e = Ά 0.75 h + 0.5 l , 1.5 Յ l / h Յ 3.5 · ͑ 1 ͒ 2.5 h , l / h Ն 3.5 These equations may result in effective flange widths greater than 1.5 times the height or three quarters of the wall height on either side of the web of the wall and will result in a significant increase in the flexural resistance of the nonrectangular masonry element. Since the new MSJC standard uses elastic design for uncracked unreinforced masonry ͑ URM ͒ ͑ Fig. 2 ͒ , the inherent upper limit on flexural capacity imposed by these assumptions will usually result in noncritical levels of shear stress at the flange-web interface. This may not be the case if a cracked section analysis is performed. Only a limited amount of research has been carried out in this area. As a consequence, code design rules vary considerably from country to country and reflect the limited knowledge available. In-depth studies of this phenomenon have been carried out in relation to diaphragm wall behavior by Phipps in the United Kingdom ͑ Phipps and Montague 1986 ͒ , and this work was ...
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Citations
... Various experimental campaigns conducted both at the scale of single masonry panels [10][11][12][13][14][15] and at the structural level [16][17][18][19] highlighted potential issues associated with the "flange effect". In addition, various numerical investigations based on refined FE models highlighted the relevance of the topic [20,21]. However, these works have investigated the issue more at the scale of single components than from the perspective of systematically assessing the repercussions at a global scale or for deriving simplified modeling strategies to be implemented in EF models, as instead is the main goal of this paper. ...
This paper focuses on the so-called “flange effect” in unreinforced masonry buildings when the connection among walls is good, thus forming a 3D assembly of intersecting piers (with L-, C-, T-, or I-shaped cross-sections). Given the direction of the horizontal seismic action, the presence of such flanges (the piers loaded out-of-plane) can influence the response of the in-plane loaded pier (the web) in terms of failure modes, maximum strength, and displacement capacity. Specific rules are proposed in codes to evaluate the effective width of the flange, for the in-plane verification of a single masonry wall. However, in the case of 3D equivalent frame (EF) modeling of the whole building, all the intersecting piers should be considered entirely, to model the response in both the orthogonal directions as well as the torsional behavior, but this may lead to overestimating the flange effect if a perfect connection is assumed. This paper investigates the capability of simulating the actual behavior in EF models by introducing an elastic shear connection at the intersection between two piers using an “equivalent beam”, coupling the nodes at the top of piers. A practice-oriented analytical formulation is proposed to calibrate such a flange effect on the basis of the geometric features and material properties of the web and the flange. Its reliability is tested at the scale of simple 3D assemblies and entire buildings as well. Finite element parametric analyses on masonry panels with symmetrical I- and T-shaped cross-sections have been performed to investigate the axial load redistribution between the flanges and the web and the consequent repercussion on the overall performance of the web. The results have proven that, after a calibration of the shear connection, the variation of axial force between the web and the flanges is correctly reproduced and the strength criteria for 2D panels provide reliable results. Finally, in the conclusions, some practical hints for simulating an imperfect wall-to-wall connection are also provided, since this case is relevant in historic masonry buildings, which are characterized by different masonry types, transformations over time, and already-cracked conditions.
... To broaden the scope of applications of structural masonry, improvement in shear strength is required and provisions must be introduced in the codes to allow for the effects of bonding patterns on the shear resistance of masonry. Experimental studies in this area were developed by Lissel, Shrive and Page (2000), Capuzzo Neto, Corrêa and Ramalho (2008), Drysdale, ElDakhakhni and Kolodziejski (2008), andBosiljkov et al. (2010). According to the authors, the webflange bonding pattern affects the interface shear strength of interconnected masonry walls and the type of connection influences their behavior. ...
... This value was obtained from previous numerical simulations (OLIVEIRA, 2014). According toBosiljkov et al. (2010), the level of pre-compression on the flanges up to approximately 0.5 MPa influences the shear strength, for higher values, the results stabilize. Afterwards, a shear load was applied monotonically to the top web of the sections under displacement control to produce shear in the flange-web interface. ...
Predicting the behavior of interconnected masonry walls is a challenging issue, given the influence of a wide range of factors, such as the mechanical properties of the materials (blocks and mortar) and the way the walls are connected to each other. In this paper, experimental results in H-shaped walls subjected to shear at the vertical interface are introduced with a numerical representation. Concrete blocks and two types of connections (running bond and U-steel anchors) were considered in the tests. Computational modelling was carried out using the Diana® FEM software to complete the study and understand the structural behavior of the masonry panels. The influence of the bonding pattern on the experimental and numerical response was studied and good agreement between the results was found. Moreover, the numerical analysis showed that the computer models of the interconnected walls adequately represented the behavior of the physical models regarding load capacity and cracking patterns.
In unreinforced masonry buildings, load bearing walls are not of only rectangular shape in the plan, but the cross walls connected to them make the walls of either I shape, T shape, C shape, etc. The presence of the cross walls can significantly change the behavior of the masonry walls in terms of strength and the failure modes of the walls. In this paper, a parametric study has been performed to investigate the effect of different parameters, namely pre-compressions, material properties, aspect ratios, and boundary conditions on the in-plane behavior of the masonry walls. The masonry walls have been modeled using the simplified micro-modeling approach in the standard finite element software Abaqus without any user-defined subroutine. It has been found that the strength and damage pattern of the flanged walls are different from the rectangular walls, which needs to be considered in the analysis of masonry walls using a numerical approach or while conducting experiments.
The vertical load distribution between the walls in a masonry building is a difficult issue in structural engineering. The absence of research regarding this subject results in using simple load distribution methods by the designers. This paper presents a study of the vertical load distribution between the walls in a full-scale four-story masonry building. Ninety blocks of a four-story building were instrumented with strain gauges and their behaviors were evaluated after an external loading was applied on some slabs. The distribution of the vertical loads in this building is also evaluated using a simple theoretical design model widely used in structural design practice. Experimental results exhibited vertical load redistribution between the walls and high influence of the flexure of slabs and walls on this distribution. Comparing experimental and theoretical results, there was similarity between the experimental and numerical results inside some wall groups, although the simple design model produced divergent stress values in the presence of significant flexural effects.
The mechanical behaviour of unbonded post-tensioned (UPT) shear walls of high-strength CAlcium SIlicate ELement masonry with Thin-Layer Mortar (CASIEL-TLM masonry) was investigated experimentally and numerically. Eight walls were tested with the following key variables: unit type, prestress level and cross-sectional shape (rectangular or T-shaped with interlocking of web and flange). An extensive measurement scheme was adopted that allowed derivation of average curvatures and strains in the bottom region of the wall in addition to wall displacements. Since UPT masonry is characterised by the absence of local compatibility between masonry and the UPT tendons, a numerical model for quasi-static, monotonic push-over analysis was developed that provides an iterative solution for the global interaction between masonry and UPT tendons. A common masonry stress–strain diagram was adopted in the numerical model. A peculiarity of CASIEL-TLM masonry is the kicker course, which reduces the stiffness of the bottom region of the shear wall. This layer was modelled with no-tension, linear-elastic behaviour and a reduced stiffness. Nevertheless, the model underestimates the experimental deformations of the rectangular shear walls, while the strength is in reasonable agreement. The walls with T-shaped cross-section failed prematurely by shear of the web-flange interface, resulting in diagonal splitting cracks in the interlocking units. This paper deals with the experimental results of UPT CASIEL-TLM masonry shear walls with rectangular and T-shaped cross-section and with the numerical modelling of the overturning behaviour of UPT shear walls with rectangular cross-section.