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Engineering Structures

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Seismic behavior and improvement of autoclaved aerated concrete infillwalls

Baris Binicia, Erdem Canbaya, Alper Aldemirb,⁎, Ismail Ozan Demirela, Ugur Uzganc,Zafer Eryurtluc, Koray Bulbulc, Ahmet Yakuta

aMiddle East Technical University, Department of Civil Engineering, Ankara, TurkeybHacettepe University, Department of Civil Engineering, Ankara, Turkeyc AKG Research and Development Division, Izmir, Turkey

A R T I C L E I N F O

Keywords:Infill wallAutoclaved aerated concreteSeismic performanceOut-of-planeIn-plane responseCyclic testExperimental behaviour

A B S T R A C T

Performance of infill walls in reinforced concrete (RC) frames is generally questionable under the combinedaction of in-plane and out-of-plane seismic demands. Despite the vast number of tests investigating the behaviorof brick masonry infill walls in RC frames, past research is limited for infill walls made of Autoclaved AeratedConcrete (AAC) blocks. In the first part of the study, six single-bay single-story half-scaled RC frames were testedunder the action of in-plane cyclic displacement excursions and/or out-of-plane pressure. It was found that thepresence of out-of-plane pressure can significantly reduce the deformation capacity of RC framed AAC infillwalls. Afterwards, a new innovative system is proposed to protect the infill walls from in-plane induced forceswhile providing out-of-plane support. Sliding connectors placed inside the AAC blocks were used to isolate theinfill wall from earthquake-induced frame deformations, while providing out-of-plane support. In order to testthe beneficial effect of the proposed system, three additional tests were conducted. The proposed systemmanaged to reach in-plane deformation demand of 2% without cracking and preserving its out-of-plane stability.

1. Introduction

Infill walls are non-structural elements used to separate archi-tectural spaces and to provide heat, sound insulation and fire resistance.The poor seismic performance of infill walls in recent earthquakes (forexample Van 2011, L’Aquila 2009) have demonstrated the need toproperly design and construct these elements. Two key problems re-garding the seismic performance of infill walls are related to the in-plane and out-of-plane vulnerabilities. Infill walls, when built adjacentto the frame elements, are subjected to compressive stresses along theirdiagonal directions, while the opposite diagonal tends to separate fromthe frame [1–3]. Under increasing cyclic lateral demands, the walls cancrack or crush at distressed locations resulting in a much softer andweaker story [4–6]. Furthermore, the floor accelerations acting in theout-of-plane direction of walls may cause instability especially whenboth sides of the wall loose contact due to in-plane deformations [7,8].Hence, mitigating the vulnerable nature of the infill walls under multi-directional loading is extremely important.

A number of different schemes have been proposed in recent years

to provide safety, deformability and energy dissipation capacity to theinfill walls. Some researchers devised an isolator steel sub-frame ele-ment with sacrificial fuses that can protect the brick masonry infillwalls while sustaining damage at the wall-frame interfaces [9]. An al-ternative method to partition the infill walls by placing several steelsheets along the height of the infill walls was proposed by Preti et al.[10]. In their method, the steel sheets between masonry partitionshelped to reduce the friction and promoted the sliding of the units inthe infill wall. This intervention prevented the wall damage due to theformation of a strut mechanism. Therefore, damage was concentratedalong a number of horizontal masonry bed joints around steel sheetsproviding a successful method of partition. An infill tie system to limitthe damage at the bed joints and provide out-of-plane support throughthe steel plates placed along bed joints was also proposed as an alter-native system [11]. Seismic brick isolators placed with different ar-rangements in an infill wall were also used to enhance the performanceof infill walls [12]. These special bricks were made of a deformablematrix with an elastic modulus being a few orders of magnitude smallerthan commercial bricks and allowed movements in one direction. Fiber

https://doi.org/10.1016/j.engstruct.2019.05.032Received 29 September 2018; Received in revised form 7 April 2019; Accepted 12 May 2019

⁎ Corresponding author.E-mail addresses: binici@metu.edu.tr (B. Binici), ecanbay@metu.edu.tr (E. Canbay), alperaldemir@hacettepe.edu.tr (A. Aldemir),

ozan.demirel@bilkent.edu.tr (I.O. Demirel), ugur.uzgan@akg-gazbeton.com (U. Uzgan), zafer.eryurtlu@akg-gazbeton.com (Z. Eryurtlu),koray.bulbul@akg-gazbeton.com (K. Bulbul), ayakut@metu.edu.tr (A. Yakut).

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T

mesh and engineered cementitious composite overlays were tried toenhance the deformability of the infill wall frame systems [13]. Theidea of employing interlocking bricks to obtain a seismically protectedinfill wall were also utilized by different researchers to improve thebehavior [14,15]. Alternatively, thermoformed recycled plastic jointswere used instead of traditional mortar joints in order to allow rigidbody movement of infill wall panel [16]. All of these developmentswere made by using hollow or solid clay brick masonry and no researcheffort was placed to improve the seismic behavior of Autoclaved Aer-ated Concrete (AAC) infill walls.

AAC is a popular choice of infill wall material due to its advantagessuch as being lightweight, providing good insulation, fire-proofing andhigh durability. Currently, AAC has about 22% of the infill wall marketin Turkey. Only recently, few studies were conducted to investigate theseismic response of AAC infilled RC frames. Researchers [17,18] testedAAC infilled RC frames under in-plane cyclic displacement excursionsto compare the behavior with brick masonry infilled RC frames. Inanother study, nine half-scaled AAC infilled RC frames under low cyclein-plane loads were tested [19]. The key test parameters were thepresence of openings, wall aspect ratio, effect of prestressing and con-struction sequence of the infill wall and the RC frame. In a recent study,the in-plane pseudo dynamic test and simulation results of a three-baythree-story half-scaled frame with AAC infill walls in its middle-baysubjected to three consecutive ground motions were presented [20]. Allof these studies focused only on the in-plane seismic response of AACinfilled RC frames and none of the studies presented any method toenhance the seismic performance of AAC infilled frames.

The objectives of the current study are: (i) to investigate the in-plane behavior of AAC infilled RC frames constructed with or withoutmesh reinforced plaster, (ii) to observe the influence of the out-of-planeloads on the in-plane deformation capacity, and (iii) to devise and test aseismic isolation system for AAC infill walls that is capable of elim-inating the damage due to in-plane demands while being stable underout-of-plane demands. The outcomes of this study are believed to pro-vide a sound alternative for constructing earthquake-safe AAC infillwalls.

2. Tests on RC frames with AAC infill walls

In the first part of the study, six specimens were tested to examinethe performance of AAC infilled RC frames. The test parameters were:(i) Presence of the gap between the infill wall and the frame, (ii) pre-sence of mesh reinforced mortar overlay, (iii) application of in-planeand/or out-of-plane loading. The details of test specimens are presentedin Table 1. Single-bay and single-story half-scaled portal frame speci-mens with planar dimensions of 2500× 1500mm were used for thetest frames (Fig. 1a). The utilized infill wall details for each specimenare also presented in Fig. 1b–e. The tested specimen was selected fromthe first story central-internal bay of the five-story prototype RCbuilding designed according to the Turkish Earthquake Code (TEC2007[21]). The dimensions of the frame span, columns and beams representtypical values in RC frames [22]. The test specimen dimensions andreinforcement details were obtained from the design of an RC prototypebuilding designed according to the Turkish Earthquake Code(TEC2007). First story central-internal bay of the prototype buildingwas selected as the prototype test model. Hence the bottom beam tiedto the strong floor represents the foundation of the prototype building.The full details of the building design can be found in [23]. Such aprocedure is a standard well-known approach in designing test speci-mens [24,25]. Due to the facility limitations, specimens were scaled by½. Such scaled specimens were used successfully in the past for thecyclic testing of frames by others [24–26]. The scale effect was mini-mized by performing dimensional analysis during the design stage ofthe test specimens. For this purpose, the stress demands on the scaledspecimen was matched to the stresses on the prototype model (i.e. fullscale model). The applied vertical loading on the columns were Ta

ble1

Summaryof

test

specim

ens.

Specim

ens

Wall-frameinterface

Infillov

erlay

Load

ing

Con

cretestreng

th(M

Pa)

Com

pressive

streng

thof

AAC

bloc

ks(M

Pa)

Com

pressive

streng

thof

Plaster

(MPa

)Te

nsile

streng

thof

fibe

rmesh

(N/m

m)

S1Non

eNon

eIn

Plan

e27

.85

––

–S2

Noga

pNon

eIn

Plan

e32

.35

2.45

––

S320

mm

gapun

derthe

beam

Non

eIn

Plan

e24

.87

2.45

––

S420

mm

gapun

derthe

beam

PlasticFibe

rMesh+

Plaster

InPlan

e24

.92.45

1.05

40

S520

mm

gapun

derthe

beam

PlasticFibe

rMesh+

Plaster

Increasing

outof

plan

epressure

28.2

2.45

1.10

40

S620

mm

gapun

derthe

beam

PlasticFibe

rMesh+

Plaster

InPlan

e+

Con

stan

tOut

ofPlan

epressure

28.2

2.45

1.15

40

S7Slider

Con

nections

PlasticFibe

rMesh+

Plaster

InPlan

e26

.92.45

0.95

40S8

Slider

Con

nections

PlasticFibe

rMesh+

Plaster

Increasing

Out

ofPlan

ePressure

26.9

2.45

1.03

40S9

Slider

Con

nections

PlasticFibe

rMesh+

Plaster

InPlan

e+

Con

stan

tOut

ofPlan

ePressure

26.9

2.45

1.08

40

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adjusted to create similar vertical stresses both for the scaled specimenand the prototype model. In addition, the diameter of all reinforcementswere scaled by ½ to obtain similar reinforcement stresses. The plainbars were used as strirrups as no deformed bars with a diameter of6mm were available in the construction market. Moreover, the ag-gregate sizes were scaled down to a maximum size of 12mm in order toconsider the scale effects in the fracture process zone [27,28]. The porestructure of AAC was not altered to consider the scaling effect due to theproduction limitations.

The details of the reinforcement used in the test specimens areshown in Fig. 1a. The column dimensions were 200× 200mm witheight 8-mm-diameter deformed bars used as longitudinal reinforcementand 6-mm-diameter plain bars used as stirrups. All the longitudinalreinforcing bars were well-anchored by welding steel plates at theirends as shown in Fig. 1a. The yield and ultimate strength of the 8mm(6mm) bars were 424MPa (448MPa) and 567MPa (504MPa), re-spectively. Stirrup spacing in the columns was 50mm at the ends ofcolumns for a distance of 300mm whereas it was 100mm in the middle

portion of columns. The presence of the slab was also considered byconstructing a flanged beam with an effective slab width of 1000mmand a slab thickness of 70mm. The total beam depth was 150mm andthe web thickness was 200mm. At the ends of the beam, confined zoneswere formed by reducing the transverse reinforcement spacing to50mm for a distance of 400mm, whereas the spacing of the transversereinforcement was 100mm for the remaining portion of the beam. Allthe stirrups of the columns and beams were anchored to the core using135 °Â hooks to simulate code-compliant detailing.

Concrete used in the test frames had a maximum aggregate size of12mm with a target cylindrical compressive strength of 25MPa. Theconcrete compressive strength was obtained by testing150mm×300mm cylinders throughout the study. Uniaxial compres-sive strength of the test specimens are given in Table 1. The unexpectedvariation of concrete compressive strength from specimen to specimen,if there is any, was compensated by adjusting the axial load to achievesimilar axial load ratios (i.e. axial load divided by the column com-pressive strength) on columns. The moment-curvature analyses

Fig. 1. Reinforcement details of the RC frame and drawings of test specimens. *Clear cover is 20mm at all edges. **This detail is supplied in Specimens S7–S9.

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conducted on the beams and columns of the test specimens revealedthat the variations in the actual compressive strength of the framescould have minimal impact on the overall load-deformation response ofthe members. In addition, all the specimens except Specimen S1 hadAAC infill walls constructed by using AAC blocks with a size of300×125×100mm, a dry density of 350 kg/m3, elasticity modulus

of 1750MPa and a uniaxial compressive strength of 2.5 MPa. The ma-terial properties of AAC blocks were determined from laboratory tests.During the course of this study, compression tests for AAC blocks wereconducted by utilizing the relevant ASTM standards (ASTM C1692-11[29] and ASTM C1693-11 [30]). In this study, the mean compressivestrength values are reported. The moisture content was 5% and theelapsed time interval between the cast of the frame and test executionwas one week and it was approximately the same for all specimens.AAC blocks in the construction market have variable thicknesses ran-ging from 100mm to 350mm with constant length and height of 600and 250mm, respectively. The AAC blocks used in this study had di-mensions of 300×125×100mm (L×h×w) following the ½ scalingof the dimensions. The design of the prototype building resulted in anaxial load ratio (axial load divided by column gross area and concreteuniaxial compressive strength) of 0.175. Accordingly, concentratedaxial forces with a magnitude of about 200 kN (depending on theconcrete compressive strength at the test day) were applied on the

Fig. 2. Test setup and instrumentation.

Fig. 3. In-plane displacement history.

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columns of each specimen to simulate axial forces from upper stories. Adistributed load of 10.25 kN/m was placed on the slabs by using steelblocks to simulate the dead plus reduced live loads.

The test setup is shown in Fig. 2. The lateral load was applied by adisplacement-controlled hydraulic actuator. Gravity-induced axialforces on the columns were simulated by using two force-controlledhydraulic cylinders acting on the columns. These hydraulic cylinderswere pin-connected to the rigid exterior frame and the column ends.Out-of-plane pressure was applied through an airbag attached to thestiffened wooden plank resting on a linear guide which was approachedto the infill wall by means of a hydraulic cylinder. The airbag systempushed the infill wall in the out-of-plane direction to the desired forceby using an air pump. Pressure inside the airbag was applied in a load-controlled manner. The pressure was maintained throughout the testsuntil a significant increase in the out-of-plane deflections or a suddendecrease in the pressure level was observed. A 2-mm-thick Teflon sheet(coefficient of friction=0.10) was placed in between the infill wall andthe airbag to eliminate friction forces (Fig. 2).

For all specimens, applied forces in the lateral and axial directionswere monitored with load cells. The story displacement was measuredby using Liner Variable Differential Transformers (LVDTs) placed at thecentroid of the slab. For specimens tested under in-plane demands,

column end deformations within the potential plastic hinge (i.e. within200mm from the column and beam ends) were measured by usingLVDTs as shown in Fig. 2a. The transducers to measure column andbeam end curvatures were placed with a gauge length equal to thecolumn (or beam) depth following the proposal of Bayrak [31]. Thedamage was observed to take place within this zone at the end of thetest of the bare frame supporting this assumption (Fig. 6). Out-of-planedeformation measurements were taken at locations shown in Fig. 2buntil the out-of-plane collapse of the wall.

The applied in-plane cyclic displacement history is shown in Fig. 3.Inter-story drift deformations (i.e. top lateral displacement, Δ1 dividedby story height H) with 0.5% increments were applied twice at everydrift level prior to the next increment. For the pure out-of-planeloading, the airbag pressure was increased up until the collapse of theinfill wall. In the case of combined in-plane and out-of-plane loading(i.e. Specimens S6 and S9), the out-of-plane pressure was held constantat 1/3 of the out-of-plane capacity of the AAC infill wall obtained fromthe test results of Specimens S5 and S8. In-plane testing was conductedwith the same displacement history given in Fig. 3 as described in[32–34].

Fig. 4. Construction details of Specimens S2, S3 and S4: (a) Mortar application for the bottom layer, (b) Insertion of an L-shaped profile, (c) Photo of AAC Infill Wallbefore plaster application, (d) Application of Foam to fill the Gap, (e) Application of the first layer of plaster, (f) Placement of fiber mesh and (g) Application of thesecond layer of plaster.

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3. In-plane behavior (Specimens S1, S2, S3 and S4)

Specimen S1 was a bare frame whereas AAC blocks were used as theinfill wall of the RC frames for Specimens S2 to S4. During the con-struction of Specimen S2, the AAC block were placed adjacent to bothcolumns and beam of the reinforced concrete frame. On the other hand,20mm gap was left between the top of the AAC infill wall and thereinforced concrete beam for Specimens S3 and S4. This gap was thenfilled with expandable foam to satisfy insulation requirements. Inpractice, a gap of 20–40mm is usually left between the AAC infill wallsand the beam in order to protect the infill wall from service level in-duced deflections of the beam. An insulation foam is later used to fillthe gap and plaster is applied over the foam for fire protection. Underearthquake induced deformations, the gap tends to close on the com-pressed regions followed by the formation of the compression strut. Thethickness of the gap determines the lateral deformation level at whichstrut action is formed. Furthermore, a smaller kinematic locking me-chanism occurs for larger gap size, since the foaming agent is usuallysignificantly softer than AAC. In this study, a 20-mm gap was employedto test the upper bound of the gap size resulting in the least amount ofkinematic locking on the upper boundary. Among these specimens, onlySpecimen S4 had fiber-reinforced mesh and plaster covering both facesof the AAC. The mesh employed for the walls was Styrol ButadienCopolymer (SBC) with a unit weight of 160 gr/m2 and grid size of 5mmin both directions. Each filament had a width and thickness of 1.2 mmand 0.5mm. The uniaxial tensile strength of the mesh obtained fromthe average of three uniaxial tensile strength tests was 31MPa based ongross area (i.e. mesh width and thickness). The details of Specimens S2,S3 and S4 are shown schematically in Fig. 1c–e. The construction se-quence of the infill walls are shown in Fig. 4 for these specimens. Itshould be noted that the usual construction practice of AAC employsattaching two L-shaped profiles after the 3rd and 6th block along the

wall height. This practice was followed as shown in Fig. 4b for Speci-mens 2–6. Following the construction of the AAC infill wall, a 5-mm-thick plaster was applied. Then, the fiber mesh grid having a gridspacing of 50mm was placed on the plaster. The second layer of theplaster with the same thickness as the first layer was applied on top ofthe fiber mesh. The application procedure is summarized in Fig. 4e–g.The material properties of the plaster and the fiber mesh are given inTable 1. The average uniaxial compressive strength of the interfacemortar was 8.4 MPa with a standard deviation of 0.98MPa. Whileconstructing AAC walls, a higher strength cementitious mortar isusually used in practice compared to the mortar commonly used inbrick masonry. In this way, the weaker chain in the AAC walls is usuallythe AAC blocks rather than the interface joints. However, the thicknessof the joints is much smaller in the AAC walls compared to the brickwalls allowing for a better alignment and plumb line. The mortar in theinterface joints used in the experimental program was obtained from anAAC producer and it was same as the product used in practice. In thisway, the common practice of application was reflected in the test spe-cimens. The interface mortar was applied through the full width of theAAC block and the vertical interface with a thickness of nearly 1mm.There were full contact between the AAC blocks and the columns(Fig. 4) at the wall frame interfaces.

The lateral load-displacement responses of Specimens S1 to S4 arepresented in Fig. 5. The ultimate drift ratio (DRu: drift ratio corre-sponding to 20% capacity drop), yield drift ratio (DRy: drift ratio de-termined by using the bilinearization procedure in FEMA356 [35]) andthe displacement ductility (DRu/DRy) measured during the tests areshown on each graph. The measured maximum column curvatures atthe ultimate drift ratio are also noted in each graph. The plastic hingingsequence of the member ends (as presented on each load deformationresponse curves) was estimated by comparing the measured curvaturewith the yield curvature for each member end. The yield curvature for

Fig. 5. Measured load-deformation responses of Specimens SI, S2, S3 and S4.

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the sections were obtained based on the section analysis of members.The damage pictures at different drift levels are given in Fig. 6 for thesespecimens. The ultimate load carrying capacity of Specimen S1 was73.6 kN and 72.1 kN for the positive and negative loading directions,respectively (Fig. 5a). The ultimate lateral load carrying capacity wasreached at about 1.3% drift ratio and this capacity was maintained untilabout 3% drift ratio. Beyond this drift ratio bar buckling of longitudinalcolumn reinforcement was observed at the column base plastic hingeregions. This specimen experienced a ductile failure mode with an es-timated ductility of about 6.

Specimen S2 experienced first cracking of the AAC wall at 0.35%drift ratio. Significant reduction of the initial stiffness occurred at about0.67% drift ratio. Strength degradation did not take place until 3%interstory drift ratio at which ultimate lateral load of about 115 kN wasattained (Fig. 5b). It can be stated that the AAC infill wall increased thestiffness and the strength of the bare frame while allowing deform-ability up to about 3.2% drift ratio. Specimen S2 had a displacementductility value of 4.6 and 6.0 for positive and negative direction ofloadings.

Specimen S3 showed a similar response to that observed in

Specimen S2, i.e. initial cracking of infill wall and reduction in stiffnessoccurred at similar drift ratios for Specimen S2 and S3. The most im-portant difference was the fact that ultimate lateral strength (approxi-mately 130 kN) occurred at about 2% drift ratio (Fig. 5c). The reductionin the drift capacity in this specimen can be attributed to the extensivediagonal cracking observed at about 1% drift ratio. The extensive di-agonal cracking may be attributed to the 20mm gap between the walland beam. In Specimen S2 without gap, the surrounding frame confinesthe infill wall and keeps crack width limited as compared to SpecimenS3. Specimen S3 had a small decrease in its ductility capacity withrespect to Specimen S2, which could be due to the presence of the gapbetween the beam and the AAC infill.

Specimen S4 behaved in a stiffer manner owing to the presence of aplaster layer with a compressive strength of 1MPa. No cracking exceptalong the interfaces was observed in this specimen up until 1% driftratio, beyond which strength degradation initiated. The ultimatestrength of about 115 kN (−135 kN) dropped to its 80% at about 2%(−2.5%) drift ratio (Fig. 5d). The displacement ductility of this spe-cimen was about 4.0 (5.0), stating a reduction in deformability com-pared to the bare frame. However, Specimen S4 had a superior wall

Fig. 6. Pictures of damaged Specimens SI, S2, S3 and S4. *Dashed lines indicate the cracks due to positive drift (rightward direction) and solid lines indicate thecracks due to negative drift (leftward direction) and hatched areas indicate the crushing.

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cracking resistance compared to those of Specimens S2 and S3throughout the test evidenced by the observed damage (Fig. 6).

Similar to the results presented above, the significant influence ofAAC infill walls on the frame behavior was previously observed byothers [36–39]. The effect of infill-frame interaction was found to bebeneficial at low lateral displacement demands improving both thestrength and lateral stiffness. However, disintegration of the AAC infillwall from the RC frame was observed at moderate-to-high lateral

displacement demands. In addition, the limited number of test resultspresented herein suggest that in the absence of plaster overlay withfiber reinforced mesh, AAC infill walls can be extremely susceptible tocracking at low deformation demands under in-plane loads. Whenplaster is used along with a fiber mesh, premature cracking diminishesalong with the reduction in the drift capacity is observed. The damagepictures shown in Fig. 6 show that the column ends of the bare frameexperienced significant inelastic action in the form of cover spalling

Fig. 7. Out-of-plane loading test results (Specimen S5).

Fig. 8. Effect of out of plane loading on in-plane deformability: (a) Comparison of load - drift ratio curves of Specimen S4 and S6, (b) OOP and IP interaction diagramand (c) Photographs of observed damages of Specimen S6 corresponding to drift ratios of 0.35%, 0.50% and 1.0%.

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initiating at about 2% interstory drift ratio and rebar buckling at about4% drift ratio. Specimens S2 and S3 without plaster and fiber meshsustained cracking starting from 0.5% interstory drift ratio. Duringcyclic displacement excursions diagonal cracking occurred along thestrut directions. The presence of the gap, despite having some influenceon the cracking sequence, did not significantly change the response ofthe infill walls. However, the presence of the plaster with fiber re-inforced mesh was extremely influential in covering the AAC cracking,if there is any. Hence from the point of serviceability the presence of asoft plaster with mesh appears to be beneficial. However, the reductionof the drift capacity of the specimen with the plastered infill wall is themost notable observation from this test.

4. In-plane and out-of-plane tests (Specimens S5 and S6)

Specimen S5 was tested under increasing out-of-plane pressure. Theobtained total out-of-plane load versus midspan deflection is shown inFig. 7 along with its pictures of observed damages. It can be observedthat the stability of the wall was lost at about 50 kN and 17mm mid-span deflection. The observed crack pattern during the out-of-plane testof Specimen S5 shows that the top line of the infill wall tends to deflectfreely due to the presence of the gap while the sides and the bottom aresupported. A simplified yield line pattern conforming to the observedcracking is shown in Fig. 7. The test results support the recently pos-tulated yield line mechanism of Vaculic and Griffith [40] to explain theout-of-plane failure of masonry walls with different degrees of top re-straints. The cracking pattern resembled a two-way mechanism, whilethe failure initiated by the loss of contact between the infill wall and

Fig. 9. The proposed system.

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beam. The out-of-plane load carrying and deformation capacities cor-responded to about twelve times the weight of the wall and 1.3% of thewall height, respectively.

In order to establish the interaction between the drift capacity andthe out-of-plane pressure, Specimen S6 was tested under constant out-of-plane pressure (i.e. total load of 17 kN which is 1/3 of the out-of-plane strength of Specimen S6) and under increasing cyclic in-planedeformations. The response of this specimen in comparison to SpecimenS4 (similar specimen tested in the absence of out-of-plane pressure) ispresented in Fig. 8 along with the observed damage. The specimenexperienced cracking while in-plane drift demand was at 0.35%. Thecracking of the wall was observed in an accelerated manner comparedto that observed in Specimen S4 and the out-of-plane collapse occurredduring the second cycle of 1% interstory drift ratio. Interestingly, thecollapse occurred while the lateral displacement was about zero as thecontact on both sides of the frame and the AAC infill interface was lost.This result shows that out-of-plane demand tends to accelerate thedamage of the compression struts during in-plane deformations and theout-of-plane collapse tend to occur when the least amount of frictionalresistance is available at the frame-infill interface. Fig. 8b presents theobtained interaction diagram in terms of in-plane (IP) Drift Ratio andout-of-plane (OOP) Load Capacity Ratio (i.e. the ratio of the out-of-plane load applied during the test to the out-of-plane load capacity ofthe AAC infill determined from the out-plane only test) from testing ofthe specimens. One can observe the adverse effect of out-of-plane

loading on the in-plane drift capacity based on the results of these twotests as the interaction appears to be linear. In other words, in thepresence of combined in-plane and out-of-plane loadings, RC frameswith AAC infill walls are not expected to satisfy the usually accepteddrift limit of 2% (i.e. TEC2007 [21]).

5. Proposed system (Specimens S7, S8 and S9)

Infill walls are non-structural elements constructed to provide in-sulation, separation, fire-proofing etc. Hence, they are not expected toalter the lateral load response of the framing systems designed bystructural engineers under seismic loads. Test results of SpecimensS1–S6 show that classical AAC infill walls in reinforced concrete framesare not expected to perform satisfactorily under future earthquakes.They are expected to crack extensively not realizing serviceability re-quirements under service level earthquakes or fail under combined in-plane and out-of-plane seismic demands. Modern seismic codes (such asthe Turkish Earthquake Code 2018 (TEC2018) [41]) require non-structural elements to be isolated from the framing systems such thatthe framing system is sufficiently ductile under lateral forces. Hence,AAC infill wall systems are in need of a protective system that: (1)Allows the RC frame system to behave as deformable as a bare frame,(2) sustains out-of-plane loads without collapse, (3) removes the ad-verse effect of out-of-plane loading in reducing the in-plane drift ca-pacity and (4) preserves the attractive features of AAC walls for

Fig. 10. (a) Assembler placed in AAC block, (b) intrusion of the Assembler, (c) vertically-placed and horizontally-placed assemblers, (d) channel connection ap-paratus, (e) the slider, (f) installment procedure of the vertically-placed assemblers and (g) installment procedure of the horizontally-placed assemblers.

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insulation and fire-proofing.To accommodate all requirements, a new innovative system was

proposed and tested. While designing and testing the proposed system,it was aimed to devise an autoclaved aerated concrete block solution forearthquake-safe AAC infill walls. The key features of the proposedsystem were chosen as the isolation of the infill wall during the in-planelateral deformations of the boundary frame and providing supportagainst out of plane loading. For this purpose, AAC blocks with theslider connectors were manufactured by inserting the metal slider as-sembly into the AAC blocks (Fig. 9). The slider was provided withsufficient allowance to move in and out of the AAC block depending onthe target design drift of the story. The proposed system was composedof assemblers placed inside the AAC block (Fig. 10a–c), the channelprofile (Fig. 10d) and the slider (Fig. 10e). The assembler was placedinside the AAC block while the AAC was in its fresh state (i.e. after heatcuring before autoclaving). The assemblers could be inserted in the AACblocks in two different ways, i.e. horizontal or vertical, according to theposition of the AAC blocks (i.e. for columns horizontally-placed as-semblers and for beams vertically-placed assemblers). The proposedsystem is just a plug-and-install type system and its installment is nottoo different than ordinary AAC infill construction. The applicationsteps of the proposed system were as follows (Figs. 9 and 10): (i) Thechannel profile was installed on the column (or beam) by using pinnails. (ii) The AAC block with the built-in slider connector was posi-tioned next to the channel. (iii) The slider end was inserted into thechannel and the block was rested on the previously laid block. Theslider was allowed to move 8 cm in and out of the block such that cyclicdeformations of up to 2% could be accommodated during the tests. Forthe beam connections, a similar procedure was used by inserting theslider into the channel connected to the beam in the horizontal direc-tion. This procedure was repeated with any desired spacing of the sli-ders selected based on the required out-of-plane strength. In this study,the spacing of the proposed system was selected as one slider per three

rows of AAC based on the preliminary tests of the slider. Each slider wasfound to carry about 3 kN out-of-plane shear and a total of eight sliderswere found to be sufficient for the desired out-of-plane capacity. Thegap at the interface of columns (beam) and the infill wall was filled withstone wool providing insulation and fire resistance (Fig. 9). The infillwall was finally covered with fiber reinforced mesh plaster overlay as afinal finish. During the installation process, the channel profiles weresecurely mounted on the RC columns and beams. On the other hand, theslider connectors were free to move in the plane of the AAC infill wallisolating it from the frame deformations. Hence, the application of theproposed system was not aimed to contribute to the lateral load re-sistance of the RC frame as can be seen from the comparison of ex-perimental results of Specimens S4-S6 and Specimens S7-S9 later in thissection.

Three specimens were tested to examine the performance of theproposed system. These specimens were replicas of Specimens S4, S5,and S6 in terms of frame properties and loading, however, the proposedsystem was installed on all three specimens. The measured load-de-flection response of the tested specimens, i.e. Specimens S7, S8, and S9,are presented in Fig. 11 along with the results of the companion spe-cimens without the proposed system. The pictures of the specimens atvarious drift levels are shown in Fig. 12. During the testing of SpecimenS7, cracking of the AAC wall was not observed. With cyclic displace-ment excursions, the gap of the connectors tended to close on theloaded wall-frame interface, whereas they opened on the opposite in-terface. Consequently, the interaction between the frame and the infillwas perfectly eliminated with the proposed system. The testing ofSpecimen S7 was stopped when the limit of drift ratio proposed byTEC2007 [21] (i.e. 2%) was attained. At this drift ratio, lateral strengthdrop was only about 8%. The ductility of Specimen S7 was about twotimes that of Specimen S4. This result shows that the proposed systemprovided sufficient ductility and deformability for in-plane loading.

The out-of-plane load carrying capacity of Specimen S8 was about

Fig. 11. Load deflection response of Specimens S7, S8 and S9.

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20 kN. Although this strength was less than that of Specimen S5, it wasstill about 6 times the weight of the AAC wall. It should be emphasizedthat the capacity of the AAC infill frames with the proposed systemcould be increased by using more sliding connectors. The testing ofSpecimen S8 revealed that the proposed system resulted in nearly 1.7times the out-of-plane ductility measured in Specimen S5. The ductilityenhancement of Specimen S8 can be attributed to the ductile nature ofthe metallic sliders with reduced yield displacement compared to thatobserved in Specimen S5. The load deformation response of SpecimenS9 exhibited only 8% strength degradation at a drift ratio of 2.5%.Consequently, the displacement ductility of Specimen S9 increased by afactor of about 2. In addition, no sudden out-of-plane collapse of theinfill wall was observed during the testing of Specimen S9. The in-planedrift capacity versus the out-of-plane load ratio of the proposed systemwas significantly improved compared to the conventional AAC infillwall system (Fig. 8b). From Fig. 8b, it could be observed that the pro-posed system can sustain up to 50% of the out-of-plane strength whileundergoing 2% inter-story drift ratio in the plane of the frame. Thisshows that while the frames with classical AAC infill walls cannot meetthe code [21,42] given drift limits, those with the proposed system aresufficiently deformable satisfying the drift limit requirements. It shouldbe noted that the scaling of the specimens by a factor of ½ results in thereduction of the out-of-plane load of ¼ [43]. Hence, the out-of-planecapacity values reported herein must be viewed in the light of thescaling factor while considering the capacity of full-scale infill walls.

6. Discussion on test results

In this section, the energy dissipation capacities and the stiffnessdegradation characteristics of the tested specimens are discussed.Firstly, the cumulative dissipated energy curves of all specimens weredetermined for the prescribed cyclic displacement excursions. Thedissipated energy for the tested specimens were calculated by in-tegrating the force displacement curves at a given drift ratio. It shouldbe noted that the dissipated energy calculations included all the cyclesat the selected drift ratio [44]. Then, the obtained cumulative dis-sipated energies were normalized by the elastic energy (i.e. base shearat the first yield× drift ratio at the first yield). The obtained normal-ized cumulative dissipated energies were plotted against the cumulativedrift ratios (Fig. 13a). It could easily be inferred from Fig. 13a that thespecimens with the proposed system (Specimen S7 and S9) had anenhanced dissipated energy capacity. In addition, Specimens S1-S4 hadvery similar energy dissipation capacities.

The stiffness degradation characteristics of all specimens were alsoinvestigated. To this end, the slope of the secant line between the originand the maximum displacement at the selected drift ratio was calcu-lated. Then, this slope was normalized by the slope of the elastic region(i.e. base shear at the first yield / drift ratio at the first yield). Thechange in the secant slope against the cumulative drift ratio was alsopresented in Fig. 13b. From Fig. 13b, it was apparent that the specimenswith the proposed system had lower stiffness values due to the reducedinteraction between the AAC infill and the RC frame. However, thestiffness degradation characteristics of all the tested specimens weresimilar to each other.

Fig. 12. Pictures of damaged Specimens S7, S8 and S9.

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The specimens with the proposed system showed a different beha-vior than their counterparts constructed with ordinary method of AACinfill wall construction. The in-plane damage of the infill wall was de-layed to larger drift ratios as the proposed system managed to separatethe AAC infill from the RC frame. More importantly, the out-of-planedamage was limited to separation cracks and some localized crushing ofthe AAC blocks at the interface of the infill and the RC frame. No totalcollapse of the AAC infill wall as opposed to the specimens without theproposed system was observed. This observation was the result of thecontrolled-slip type motion devised by the proposed system.

7. Conclusions

In this study, one bare and eight AAC-infilled frame tests wereconducted to understand the influence of (i) plaster overlay, (ii) out-of-plane and in-plane interaction, and (iii) the proposed wall system onthe response of AAC infilled frames. The following important conclu-sions can be drawn based on the results of limited number of testsconducted in this study:

(1) In the absence of a plaster layer, AAC infill walls were prone tocracking at low deformation levels (i.e. service level earthquakes).

(2) Presence of the gap between the beam and AAC infill layer was notfound to affect the infill performance significantly.

(3) The presence of fiber mesh plaster eliminated the visible crackinghence increases the serviceability at low deformation demands. Onthe other hand, drift capacity was reduced.

(4) The out-of-plane loading reduced the drift capacity almost linearlyas a function of the out-of-plane pressure. Hence, combined in-plane and out-of-plane demands should be considered in the designof AAC walls.

(5) The proposed innovative system was capable of removing theframe-infill interaction and improving combined in-plane and out-of-plane behavior (i.e. the reduction in the drift capacity due to the

presence of out-of-plane loading). Hence, it appears as a viableimprovement for AAC infill walls in seismic zones.

(6) Practical numerical simulations of the test specimens could beperformed by modeling the infill walls by using struts with in-planeand out-of-plane rigidity to consider the interaction between the in-plane and out-of-plane effects [45]. For this purpose, the strutsrepresenting the infill wall could be defined with fiber elements orlumped plasticity models. The numerical models of the specimenswith the proposed system could be generated by also defining in-terface elements between AAC infill wall and the surrounding RCframe. The experimental findings in this study should be utilized toverify the numerical models. This task is beyond the scope of thiswork and is suggested for future studies aiming to develop designand analysis methods for AAC infill walls.

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