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This paper proposes a partial-destructive method for connecting adjacent orthogonalmasonry walls by means of aramid fibre reinforced polymer (AFRP) rods, in order toimprove the overall performance of masonry buildings under horizontal forces. The proposedmethod is supported by an experimental campaign to assess the effectiveness of the

strengthening measure, and by an analytical study to develop equations suitable for design.The experimental tests showed that the connection between adjacent masonry walls isactually effective in increasing both their strength and stiffness. It is also shown that the

developed analytical equations satisfactorily predict the relevant design quantities.Keywords:connection of adjacent walls; strengthening of URM; improvement of flexural behaviour;out of plane overturning collapse; FRP strengthening.

1. INTRODUCTION

It is common to find in our cities old buildings with load bearing masonry walls that were designedonly to resist vertical loads and with no consideration of horizontal loads. This is usually reflectedin the arrangement of walls perpendicular to each other, which are not clamped along the edges.

This lack of connection may be found either in backbone (or main) walls that intersect with eachother, or in other walls considered as secondary, but made of stone that are given a structuralfunction, or even in walls that were independent in the original construction, because designed asvertical load bearers or because partly independent and without a load-bearing function.

When seismically retrofitting these buildings, it may be convenient to take advantage of all existingwalls to provide resistance to horizontal forces by connecting them to each other at their intersections.By doing so, two originally unconnected walls are made into a single one with a T-shaped cross-section. The result is that the strength and the stiffness of each single wall are increased. This allowsavoiding insertion of additional walls or thickening of the existing ones, with all the obviousdifficulties related to these strategies. In the past, such strengthening measures were carried out by

STRENGTHENING OF MASONRY WALLS BY TRANSVERSE CONNECTIONTHROUGH AFRP RODS: EXPERIMENTAL TESTS AND ANALYTICAL MODELS

NED UNIVERSITY JOURNAL OF RESEARCH, THEMATIC ISSUE ON EARTHQUAKES, 2012 61

Marco Vailati1, Giorgio Monti2

Manuscript received on 17thJune 2012, reviewed and accepted on 21stAugust 2012 as per publication policiesofNED University Journal of Research.

ABSTRACT

1Post-Doctoral Researcher, Department of Structural and Geotechnical Engineering, Sapienza University of Rome, Italy. Ph. (+39) 06 49919254,

Fax. (+39) 06 3221449, Email: [email protected] Professor, Department of Structural and Geotechnical Engineering, Sapienza University of Rome, Italy. Ph. (+39) 06 49919197,

Fax. (+39) 06 3221449, Email: [email protected].


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M. Vailati and G. Monti

NED UNIVERSITY JOURNAL OF RESEARCH, THEMATIC ISSUE ON EARTHQUAKES, 201262

Marco Vailati is a Post Doctoral Researcher in the Department of Structural Engineering andGeotechnics at Sapienza University of Rome, Italy where he received his Masters in Innovationin Design, Rehabilitation and Control of Structures: Assessment and Retrofitting in SeismicAreas and PhD in Structural Engineering, respectively, in 2004 and 2011. His research Interestsinclude earthquake risk assessment of building structures, and strengthening of masonry wallsand reinforced concrete structures with advanced techniques.

drilling holes and inserting metal bars in them, which were then partially filled with mortar injections.However, metal bars are often subjected to corrosion if not properly injected.

Therefore, a technique to connect two originally unconnected walls into a T-shaped layout has beenstudied and tested, which makes use of thin rods of aramid fibre reinforced polymer (AFRP) insertedin small holes drilled in the flange wall. These rods are then anchored to the lateral surfaces of theweb wall through spread fibres, which are then glued onto the web wall surface, thus providingefficient anchorage. Since the connection has to restrain the vertical sliding between the connected

walls, the rods are placed at 45, so as to provide tension components in both sliding verses.

In this way, the strengthening technique maximizes the capacity of the existing structural walls withminimum invasivity.

2. TESTING SET-UP

The tests described below have been carried out at the Department of Structural and GeotechnicalEngineering of the Faculty of Architecture of the Sapienza University of Rome, Italy.

2.1 Geometry and Boundary Conditions

The configuration of the testing rig with the walls is shown in Figure 1.The boundary conditions are:FV= 0 . t. l; uv1= uv2= 0; uh= imposed;Fh= measured.where 0is the average stress in the web wall cross-section; t and l are the web thickness and length,respectively; uv1and uv2are the vertical imposed displacements at the two flanges; uhis the horizontalimposed displacement;Fhis the horizontal measured force. Therefore, it can be understood that alltests were performed under displacement control, in order to follow any possible degrading branchin the wall response.

Figure 2shows a typical wall configuration, along with a detail of the connection with the anchorage.The measuring equipment is constituted by five strain gauges, numbered from 1 to 5; the first fourmeasure the vertical relative displacements between web and flange, while the last measures thehorizontal absolute displacement.

Giorgio Monti is a Full Professor at Sapienza University of Rome, Italy. His research interestsinclude modelling, analysis and assessment of reinforced concrete and masonry structures

under seismic excitation, structural health monitoring, strengthening techniques with innovativematerials (FRP), strategies for the preservation of historical towns, and reliability analysis ofstructures and infrastructures in seismic zones. He is an active member of national andinternational committees for the development of seismic design codes.

Figure 1. Configuration of testing rig and walls.


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2.2 Materials

Clay bricks with premixed mortar were used to build the walls, while AFRP rods were used toconnect them. The walls mean mechanical properties were: compressive strength (fd) = 6.00 MPa(0.87 ksi); shear strength (d) = 1.38 MPa (0.2 ksi); Youngs modulus (E) = 2,700 MPa (392 ksi);shear modulus (G) = 900 MPa (131 ksi).

The AFRP rod properties were: tensile strength (fyd,c) = 1400 MPa (203 ksi); Youngs modulus (Ed,c)= 60,000 MPa (8.7x106ksi)

The total tensile strength of the connection system was: (a) configuration 1RT,1= 32 kN (7 kips);(b) configuration 2RT,2= 16 kN (3.6 kips).

In configuration 1, the spread of the terminal anchors lays along the bar axis, while in configuration2 it is arranged at 90 (Figure 3).

2.3 Connections between web and flange walls

The connection between web and flange walls is constructed in few simple steps (Figure 4). Thedetails of these phases are as under

Phase a: drilling holes from the outer face of the flanges, tangent to the web wall facesPhase b: inserting the rods through the flanges and fixing them in the hole with mortarPhases c and d: gluing the spread ends to the surface of both wallsPhases e and f: applying aramid sheets with vinylester resin to cover and strengthen the anchorages

The AFRP rods were of 5.5 mm (0.22 in.) diameter, while the holes crossing through the flangewere of 7 mm (0.28 in.) diameter.

3. TEST RESULTS

The walls specimens consist of one web wall and two flange walls at its ends. They are instrumentedas shown in Figure 2.

Figure 2. Displacement transducers on a masonry wall specimen. At top right the anchoragedetail of the rods.

Note: All dimensions are in mm; 25.4 mm = 1 in.


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Figure 4. Construction phases of the connections between flanges and web walls.

Figure 3. Detail of AFRP rod anchorage: the two reference configurations (above), the testconfiguration (below).


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With reference to Figure 2, transducers 1, 2 and 3, 4 are used to measure the relative displacementbetween the web wall and flanges, while transducer 5 measures the horizontal displacement betweenfoundation and wall head.

The global effects of this strengthening method are illustrated in Figure 5-7. Figure 5shows thecomparison at the first cycle between unstrengthened and strengthened wall. In the strengthenedwall, the improved collaboration between the orthogonal walls gives rise to an increase in the initialstiffness. Note that, in the strengthened wall, the stiffness is increased approximately by 40%.

Figure 5. Comparison between strengthened and unstrengthened wall at the first cycle: increaseof stiffness.

Figure 6. Horizontal force vs. relative displacement between web and flanges walls.

Figure 7. Comparison between strengthened and unstrengthened wall at ultimate: increaseof flexural capacity. At top left: detail of crushing of masonry.

Note: 25.4 mm = 1 in.; 4.448 kN = 1 kips

Note: 25.4 mm = 1 in.; 4.448 kN = 1 kips

Note: 25.4 mm = 1 in.; 4.448 kN = 1 kips


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For low horizontal forces, the two curves are practically coincident; however, when passing 30%of the unstrengthened wall failure load, the connections modify the wall system response. This delaycould eventually be avoided, if need, by pretensioning the rods. In fact, a truss-like mechanism canbe activated only when the rods are sufficiently stressed in tension.

In Figure 6, the difference in terms of relative displacement between the web and flange walls canbe observed: for the unstrengthened configuration it is 0.31 mm (0.01 in.), while in the other oneit is 0.02 mm (7.8x10-4in.). The smaller displacement in the strengthened case is due to the presenceof the rods.

Initially, the contacting surfaces slip relatively one to each other, however, when the rods start beingpulled, they provide their contribution by imposing compatibility of the displacements of bothsurfaces. After this stage, the wall system shows a significant increase of strength and ductility, asshown in Figure 7.

In Figure 7, the curves were obtained by inverting the load path at yield displacement (curve 1),at 50% of ultimate displacement (curve 2), at ultimate displacement (curve 3). Note that each reversalpoint is marked by a circle.

As observed from the tests performed, the strengthening effectiveness strongly depends on the

correct application of the rods. In order to exploit the material mechanical properties, the rods areinserted at 45. As a matter of fact, for a plane stress state in shear, the maximum tensile forceis inclined at 45 with respect to the vertical, as shown in Figure 8.

Two effects can be observed when reaching the ultimate limit state

1) Detachment from the wall of the aramid sheet that covers the rods anchor;2) Loss of the flange wall verticality

In Figure 9b, a tension failure of a rod is shown, with clear signs of breakage in its cross-section.

Figure 8. Shear transfer mechanism between web wall and flange.


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Figure 9. Collapse by flexural mechanism of wall: (a) collapse of masonry with expulsion ofmaterial to external direction; (b) traction failure of a rod.

4. COMPARISON WITH ANALYTICAL MODEL

4.1 Unconnected Walls

The stiffness obtained from the experimental results is compared to the analytical model proposedby Tomazevic [3]

(1)

Figure 10shows the comparison between experimental and predicted stiffness.

It can be seen in Figure 10that Eq. (1) accurately predicts the elastic experimental stiffness, while

a 50% reduction gives a good estimate of the ultimate displacement.

Figure 10. Comparison between experimental and predicted stiffness and peak strength atfirst cycle.

G.AWK

I G h

lE

22 h1 1+

=

. . . . .q ro p

Note: 25.4 mm = 1 in.; 4.448 kN = 1 kips

(a) (b)


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The wall capacity is assessed by considering two possible failure mechanisms [1, 2] that maydevelop, i.e., flexural and diagonal shear, respectively, as follows

(2)

(3)

The prevailing collapse mechanism is the one with lower capacity.

Table 1contains the parameters used to calculate the stiffness (Eq. (1)) and the capacities (Eqs. (2)and (3)).

For the case at hand, the capacities are:fyF= 64.6 kN;fyV= 390.8 kN. Therefore, it is recognizedthat failure is of the flexural type.

The experimental test on the unstrengthened walls exactly shows this behaviour, for a horizontalforce equal to 70 kN (15.7 kips), very close to the analytically predicted value. Figure 11showsthe crack pattern, typical of a flexural failure mechanism (note that the flange walls, though present,are not connected).

4.2 Connected Walls

The connection of the web wall to the flange walls improves the performance of the overall system,by increasing its bending capacity thanks to the change in shape of the base cross-section, which is

h

mm

(ft.)

1850(6)

l

mm

(ft.)

1550(5)

t

mm

(in.)

120(5)

Aw

mm

(in.)

1.86x105

(7x103)

0.83

G

MPa

(ksi)

900(130)

E

MPa

(ksi)

2700(392)

fm,d

MPa

(ksi)

6.00(0.87)

0,d

MPa

(ksi)

1.38(0.20)

0

MPa

(ksi)

0.97(0.14)

Table 1 - Parameters used to evaluate the wall capacity

Figure 11. Crushing of masonry in web wall: left and right side, near foundations.

F

2

=

q rofy H

0

p2

110 0

0.85 fd

.

t .

.

Vfy

tb

= 1 . . ..1 5 d

. .1 5 d

10

+

w


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now T-shaped, with a significant contribution from the compressed flange. The shearing capacity,however, remains unaffected and equal to that of the web wall, since the flanges do not provide anycontribution to it.

The bending capacity of the T-shaped cross section is obtained by writing the equilibrium equationof the web and flange system, considered as fully connected. Figure 12shows a comparison of thetwo configurations for adjacent and connected walls.

The equilibrium equation for the T-shaped wall is given in Eq. (4)

V .H = Np .ep +Na .ea (4)

Rearranging Eq. (4) and introducing non-dimensional notation, the following capacity equation isobtained

(5)

where MU,Tis the flexural capacity of strengthened T-shaped wall; MU,Iis the capacity of the(unstrengthened) I-shaped wall and is given in Eq. (6); , and are non-dimensional factors (Eq.(7)). Note that the first factor is of mechanical nature whereas the others are related to geometry.

Figure 12. Notation of variables and stress distribution before failure: isolated wall (left);connected wall (right).

(6)

(7)

Applying Eq. (5) to the specific case of the test walls, characterized by the following parameters:= 0.077, = 4.58, = 0.23, it is possible to assess the magnitude of the force that triggers thebending failure mechanism (106 kN (23.8 kips)). The result is practically identical to that obtainedin the test, of 110 kN (24.7 kips).

In order to calculate the capacity of the strengthened system, in this case it is also necessary to dulyaccount for the contribution of the compressed flange. As it is known, Eq. (1) takes into accountboth bending (KB) and shear (KS) stiffness of the walls (Eq. (8)).

(8)

MU,T MU,I= . (1+

+ rq

MU,I =12t0 rq

2

0 = = =0.85.fd

; ta1p

1atp

;

KI =KB + KS =12EJ

h3 1.2h+

GA


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As mentioned above, the presence of the flange only modifies the flexural response of the wall,leaving the shear behaviour almost unaffected.

It will, therefore, be sufficient to express the second moment of area for a T-shaped section in placeof that for the I-shaped section, in order to predict the stiffness of the strengthened system. Thesecond moment of area for the T-shaped section is given as

By re-defining Eq. (8) we obtain

(10)

In Eq. (10), assumes its usual meaning. Note that the second moment of area of the T-shapedsection significantly contributes to the flange width, i.e., the part of the wall at the end involved inthe flexural resisting mechanism.

Although there is a dependency on the thickness ratios of the connected walls this type of problemrequires specific study, which can only be addressed in qualitative terms here. Considering this, itmay be helpful to address the problem according to Tomazevic [3], so that the flange width is definedas in Figure 13.

The proposed approach must be considered in the context of a more general application of themethod in real cases, for buildings made of bricks.

In the specific case of the walls tested, the problem is not so significant, since all flanges lengthswere entirely involved in the resisting mechanism.

4.3 Design of the connections

Based on the above discussion, a design equation is proposed here for an easy application of themethod. The equation is capable of correlating the sliding force between the faces of the twoconnected walls with the number of rods, which is essential for a correct design of the connectionsystem.

By looking at the distribution of contact stresses in Figure 12, the sliding force between the weband flange walls can be evaluated. Since the vertical load is only applied on the web wall, the stressat the flange base is the reaction to the combined compressive force and bending moment generated

by the horizontal force. The sliding force at the interface is thus given by

Figure 13. Definition of the geometry of the flange in the connected walls.

(11)Fs= (0.85.fd- 0).ta1a

KT=GAw

GAw

12EJT1.2h2

1

h+q r..

(9)[la.(lp+ ta)

2- l2p.(la- tp)]2- 4.la.lp.(lp+ ta).(la - tp).[(lp+ ta) - lp]

2

12.[la.(lp+ ta)- lp .(la- tp)]JT=


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The number of rods to be inserted to ensure that the connection resists until bending failure is thengiven by

(12)

where RT,1-2is the connection tensile strength, according to one of the possible configurations, asindicated in section 3. The total number of the rods to be applied is therefore equal to 4n, becausethey are applied on both faces of the web wall and because they are inserted at 45.

5. CONCLUSIONS

The comparison of the performance of connected walls with that of adjacent walls (based on theexperiments carried out) indicated that a strength increase of R= 1.57 and displacement increaseof d= 3.2 was obtained. As a result, there is an increase in the capacity of about 60%, while ductilityincreases by a factor of 3. Note that, beyond affecting the main resisting mechanisms, the proposedstrengthening technique provides both local and global ductility as this application can reverse thehierarchy between failure mechanisms (shear and flexure), thus favouring the latter, which is moreductile. Finally, it was ascertained that the developed equations are accurate in predicting both thecapacity and the stiffness of the strengthened wall system.

ACKNOWLEDGMENTS

The authors wish to thank the SACEN Company of Naples, Italy, for providing and applying theNAILTEX AFRP rods.

NOTATION

fm = Mean compression strength0 = Mean shear strengthKe = Stiffness of masonry wallAw = Shear area of I-section

h = effective height of walll = Length of wall = Factor dependent of boundary condition (0.83 or 3.33 in case of cantilever)t = Thickness of wall0 = Mean normal tension on the total section areaH0 = Point along the wall where the moment change signb = Factor depending of wall slenderness, and it can be take 1Ob=h/lO1.5fyF = Flexural capacity of wallfyV = Diagonal shear capacity of wallNP = Resultant of normal force on web sectionNa = Resultant of normal force on flange sectionep = Eccentricity of Npea = Eccentricity of Na

KI = total stiffness of wall (I-shaped section)JT = moment of inertia of T-shaped sectionKT = total stiffness wall (T-shaped section)FS = Sliding force at the interface web/flangen = Number of rods

REFERENCES

[1] Ministry of Infrastructure and Transportation of Italy. Istruzioni per lapplicazione delle nuovenorme tecniche per le costruzioni di cui al decreto ministeriale 14 gennaio 2008. Supplementoordinario n. 27 alla Gazzetta Ufficiale, 2009.

[2] Ministry of Infrastructure and Transportation of Italy. Nuove Norme Tecniche per Le Costruzioni.

Gazzetta Ufficiale della Repubblica Italiana 2008.[3] Tomazevic M. Earthquake resistant design of masonry buildings. Series on Innovation in

Structures and Construction., England, Imperial College Press London, 1981. p. 109-158.

Fsn = round

2 .w 2 . RT,1-2o p


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[4] Menegotto M, Monti G, Salvini S, Vailati M. Improvement of Transverse Connection ofMasonry Walls through AFRP Bars. In: Ye L, Yue Q, Peng F, Editors. Proceeding of 5thInternational Conference on FRP Composites in Civil Engineering. Beijing, China: 2010.p. 947-950.

[5] Anthoine A, Magonette G, Magenes G. Shear-compression Testing and Analysis of BrickMasonry Walls. In: Proceedings of 10th European Conference on Earthquake Engineering.Rotterdam, Netherlands: 1995.

[6] Magenes G, Calvi GM. Cyclic Behaviour of Brick Masonry Walls. In: Proceedings of 10thWorld Conference on Earthquake Engineering. Rotterdam, Netherlands: 1992.

[7] Magenes G, Calvi GM. In-plane Seismic Response of Brick Masonry Walls. Earthq Eng StrucDyn 1997;26(11):1091-1112.

[8] Turnek V, Cacovic F. Some Experimental Results on the Strength of Brick Masonry Walls.In: Proceedings of the 2nd International Brick Masonry Conference, Stoke-on-Trent, England:1971. p. 149-156.

[9] Turnsek V, Sheppard P. The Shear and Flexural Resistance of Masonry Walls. In: Procedureof the International Research Conference on Earthquake Engineering. Skopje, Macedonia:1980. p. 517-573.

[10] Turco V, Secondin S, Morbin A, Valluzzi M.R, Modena C. Flexural and Shear Strengtheningof Un-reinforced Masonry with FRP Bars. Comp Sci Tech 2006;66(2):289-296.

[11] Tumialan G, Micelli F, Nanni A. Strengthening of Masonry Structures with FRP Composites.

In: Chang PC, Editor. Proceedings of the 2001 Structural Congress and Exposition. WashingtonDC, USA: 2001. p. 1-8.

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