Flexural displacement response of NSM FRP retrofitted masonry walls

Construction and Building Materials 49 (2013) 1032–1040

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Flexural displacement response of NSM FRP retrofitted masonry walls Michael C. Griffith ⇑, Jaya Kashyap, M.S. Mohamed Ali School of Civil, Environmental and Mining Engineering, University of Adelaide, Adelaide, South Australia 5005, Australia

h i g h l i g h t s " This paper presents flexural test results of 15 FRP strengthened masonry walls. " Cyclic loading, pre-compression, reinforcement ratio and strip spacing were studied. " Cyclic loading and vertical pre-compression shown to have little effect. " Strip spacing and reinforcement ratio strongly affect wall performance. " Smaller strips at closer spacing gave best strength and displacement response.

a r t i c l e

i n f o

Article history: Available online 24 July 2012 Keywords: Brick masonry Earthquake FRP Flexure IC debonding Displacement capacity

a b s t r a c t Fifteen clay brick masonry wall tests were conducted to study the behaviour of near-surface-mounted (NSM) carbon fibre reinforced polymer (CFRP) retrofitted masonry walls in flexure. All walls were simply-supported along their top and bottom edges and tested in vertical one-way bending under threeand four-point bending and failed by intermediate crack debonding. The results of these experimental tests indicate that: (1) reverse cyclic loading and axial pre-compression had little effect on wall load– deflection response; (2) FRP strip spacing and reinforcement ratio strongly influenced wall response; and (3) for a constant reinforcement ratio, smaller strips at closer spacing gave better wall strength and displacement response. Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction Unreinforced masonry (URM) structures constitute both a significant portion of the world’s heritage buildings and a significant component of the modern residential building stock, and are particularly susceptible to damage from out-of-plane loads such as those generated by earthquakes (Ingham and Griffith [1]). Catastrophic out-of-plane flexural failures of URM (hereafter termed ‘masonry’) walls, parapets and chimneys during seismic events worldwide (e.g. L’Aquila, Italy in 2009 and Christchurch, New Zealand in 2010–2011) continue to highlight the need to strengthen these structures. Many of the traditional strengthening techniques available today for enhancing the structural performance of masonry such as steel plate bonding, steel frame works, shotcrete jacketing have disadvantages such as adding considerable mass to the structure, being labour intensive, creating working space and access limitations, and perhaps most significantly impinging on the aesthetics ⇑ Corresponding author. Tel.: +61 8 8303 5451; fax: +61 8 8303 4359. E-mail addresses: [email protected] (M.C. Griffith), jkashyap @civeng.adelaide.edu.au (J. Kashyap), [email protected] (M.S. Mohamed Ali). 0950-0618/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.conbuildmat.2012.06.065

of the building. Hence, the use of fibre reinforced polymers (FRPs) for seismic retrofit work of concrete and masonry structures has gained much attention (Hamed and Rabinovitch [2]; Willis et al. [3]; Petersen et al. [4]). In particular, over the past decade or so, near surface mounted (NSM) FRP has emerged as a promising technique due to the advantages it offers over externally bonded (EB) FRP (Fig. 1). The NSM FRP retrofitting technique (i.e. inserting FRP strips into grooves cut into the surface of a wall) provides significant advantages over externally bonded (EB) FRP such as improved aesthetics, reduced surface preparation and better protection from UV exposure and vandalism. Importantly, NSM FRP debonds at higher strains than EB FRP and thus leads to more efficient use of the FRP material (De Lorenzis and Teng [5]). For a wall supported on all four sides and subjected to out-ofplane bending, the vertical bending of the wall is the weakest link (Willis et al. 2010 [6]). The use of vertically oriented NSM FRP to strengthen such walls, has been shown to significantly increase the vertical bending capacity and thus the ultimate wall capacity (Korany and Drysdale [7]; Willis et al. [6]). Some of the common out-of-plane failure mechanisms of FRP strengthened masonry walls include sliding of the masonry units, flexural-shear cracking, FRP rupture, FRP debonding, punching shear and crushing of brick in compression (Albert et al. [8]; Tumialan et al. [9]; Ghobarah and

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cyclic test techniques for which the walls only had a narrow width. All further tests were conducted on 1 m wide walls to study the effects of multi-strip reinforcement. Walls 5, 6 and 8 were used to investigate the influence of strip spacing on flexural behaviour of retrofitted wall, as shown in Table 1. The effects of axial loading on the flexural response under monotonic loading was considered by tests on Walls 6, 10 and 14 while reverse cyclic loading was investigated with Walls 11, 12 and 15. The effect of cycling loading was investigated under three pre-compression conditions, i.e. with no pre-compression (Walls 6 and 12), 0.1 MPa (Walls 10 and 12) and 0.2 MPa (Walls 14 and 15). Further, the effect of reinforcement ratio was studied through Walls 4, 5, 7 and 9.

Fig. 1. FRP retrofit techniques (cross-section view).

Galal [10], Galati et al. [11]; Mosallam [12]). Among the debonding mechanisms, intermediate crack (IC) debonding governs the increase in moment capacity and sectional ductility. For that reason, the walls in this study were designed to fail by IC debonding as previous research in the reinforced concrete (RC) area (Teng et al. [13]; Oehlers et al. [14]) indicates that this failure mode is the most ductile and hence is the preferred failure mechanism for FRP strengthened flexural members. Whilst the application of NSM FRP strips appears to be a particularly viable retrofitting technique (Dizhur et al. [15]), limited research has previously been conducted on the application of this retrofitting technique to masonry structures. Further, the effect of out-of-plane cyclic loading on the load-deformation behaviour of FRP retrofitted masonry members also warrants more attention. As a result, there is a significant need for more experimental and analytical investigations on NSM FRP retrofitted masonry walls before it can be confidently used for seismic retrofit of URM Walls.

2.2. Material properties The material properties for the masonry walls and CFRP strips are given in Table 2. The masonry properties (Table 2) were determined from material tests conducted in accordance with AS3700 [16] and AS/NZS4456.15 [17]. The walls were constructed by a qualified brick layer using clay brick units and mortar consisting of Portland cement, hydrated lime and sand in a 1:1:6 ratio by volume. The CFRP material properties shown in Table 2 were obtained from the manufacturer’s data sheets although the elastic modulus was verified experimentally in separate bond-pull tests reported elsewhere (Kashyap et al. [18]). 2.3. Specimen design Each specimen was a single leaf clay masonry wall with nominal dimensions as shown in Table 1. The FRP strips were aligned vertically along the brick units avoiding the perpend joints (except for Wall 4) as it provides the most efficient increase

Table 2 Material properties for wall test specimens. Parameter

Mean

Masonry properties

(MPa)

Standard deviation (MPa)

Flexural tensile strength of the masonry, fmt Compressive strength of the masonry, fmc Lateral modulus of rupture of the brick unit, fut Elastic modulus of masonry, E Elastic modulus of brick, Eb Elastic modulus of mortar, Em

0.48

0.13

0.27

17

2.95

0.17

3.13

0.84

0.27

10700 19500 2300

2400 3700 870

0.22 0.19 0.38

2. Experimental study 2.1. Test plan Fifteen walls were tested in this study to investigate the effect of typical design variables on the performance of NSM CFRP strengthened masonry walls. The test variables included: (1) reverse cyclic loading; (2) axial pre-compression; (3) FRP strip spacing; and (4) reinforcement ratio. Table 1 shows the details of the walls tested in this research where the notation used for the walls in Column 1 indicate ‘‘Number/M or C/1, 2 or 3/O or D’’ where M or C refers to monotonic or cyclic loading; 1, 2 or 3 refers to how many NSM strips the wall had on a side; and O or D indicates whether the wall had NSM strips on one or both sides. Column 2 refers to the thickness (b) and width (d) of the FRP strip, respectively; Column 3 gives the spacing between strips (effectively equal to the wall width where one strip was used); and q (Column 4) is the reinforcement ratio. It should be noted that tests on Walls 1–4 were conducted as pilot tests to refine the monotonic static and quasi-static reverse

CFRP Properties (mean values) Elastic modulus of FRP strip, Ep Ultimate tensile strength, frupt

Coefficient of variance

165  103 MPa 2700 MPa

Table 1 Out-of-plane bending test walls. Wall reinforcement details

Test results

Wall ID (1)

bxd (mm) (2)

Spacing (mm) (3)

q¼

1M1D 2M1D 3C1D 4M1D 5M1O 6M2O 7M3O 8M3O 9M1O 10M2O 11C2D 12C2D 13C2O 14M2O 15C2D

3.6  10 3.6  10 3.6  10 3.6  10 7.2  10 4.8  7.5 3.6  10 4.8  5 3.6  10 4.2  10 4.2  10 4.2  10 4.2  10 4.2  10 4.2  10

355 355 355 230 1070 535 357 357 1070 535 535 535 535 535 535

0.092 0.092 0.092 0.142 0.061 0.061 0.092 0.061 0.031 0.071 0.071 0.071 0.071 0.071 0.071

Afrp Awall

(%) (4)

emax (5)

Fmax (kN) (6)

D at Fmax (mm) (7)

Dult (mm) (8)

Pexp (kN) (9)

emax/erupt (%) (10)

0.00983 0.00998 0.00829 0.00818 0.00699 0.01098 0.01188 0.01245 0.00947 0.01109 0.00995 0.00953 0.01117 0.01055 0.00821

18.6 18.4 +15.6/15.6 15.6 17.1 27.0 41.0 36.6 12.4 30.8 +27.3/27.6 +28.9/29.1 +30.2/1.4 29.7 +27.8/26.6

43.7 34.9 +28.8/23.1 26.5 44.0 61.0 70.3 75.0 51.8 66.2 +48.7/48.8 +48.4/48.2 +54.3/9.2 48.1 +36.3/40.0

44.3 36.6 +29.5/23.6 27.6 58.0 70.0 75.7 78.9 61.2 69.3 +53.7/64.5 +56.8/53.0 +69.4/57.0 52.0 +60.0/55.0

58.4 59.3 49.3 48.6 83.1 65.3 70.6 49.3 56.3 76.9 69.0 66.0 77.4 73.1 56.9

70.3 71.3 59.3 58.5 50.0 78.5 84.9 89.0 67.7 79.2 71.1 68.1 79.8 75.4 58.7

Notes: (i) Wall thickness t = 110 mm for all walls. (ii) Walls 1–3 have width w = 355 mm and height h = 1710 mm. (iii) Wall 4 has width w = 230 mm, height h = 1710 mm. (iv) Walls 5–15 have width w = 1070 mm and height h = 2310 mm. (v) Walls 10 and 12 have vertical pre-compression rv = 0.10 MPa, Walls 14 and 15 have rv = 0.20 MPa, all other walls have rv = 0 MPa.

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602

Support

FRP Strip

FRP Strip

258

1710

Loading point

602

Loading point

355

230

FRP Strip

110

Support

10

355

230

10

110

FRP Strip

3.6

Elevation

Cross section

(a) Walls 1, 2, 3,

3.6

Cross section

(b) Wall 4

Fig. 2. Specimen details – Wall 1 to Wall 4.

1070 110

1070 Support 535

535

Wall with one FRP strip

110

1032

1070

267

536

267

Wall with two FRP strips 2312

Loading point

110

1070

1032

178

357

357

178

Wall with three FRP strips

(b) Cross-section (Walls 5-10, 13, 14)

307

1070 456

307 110

Support 267

(a) Elevation

536

267

(c) Cross-section (Walls 11, 12, 15) Fig. 3. Specimen details – Wall 5 to Wall 15.

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Fig. 4. Out-of-plane bending test setup for Walls 5–15.

in ultimate capacity and provided the closest possible representation of an FRP strip through a homogeneous material. The FRP was obtained from the manufacturer in rolls of 1 m wide sheeting, 1.2 mm (Walls 1–9) or 1.4 mm (Walls 10–15) thick. It should be noted that the 1.4 mm thick strip was used only due to unavailability of the strip with 1.2 mm thickness, giving the closest match. All strips were fabricated by cutting and gluing the required number of individual strip elements together and spanned for the full length of the specimen. The FRP retrofitting scheme was designed using full interaction theory so that IC debonding was the failure mode, rather than tensile rupture of the FRP or crushing of the masonry. The groove for the NSM strip was cut using a diamond blade circular saw, cleaned with a high-pressure air jet, and then filled with epoxy adhesive. The strip was cleaned with acetone to remove any foreign substances before being inserted into to the epoxy-filled groove and allowed to cure for 7 days. The FRP strip was positioned flush with masonry surface for all specimens.

As shown in Table 1, dimensions of Walls 5–15 were different to that used in pilot tests (Walls 1–4) and were chosen to investigate lower reinforcement ratio and to study the influence of strip spacing. The four pilot tests (Walls 1–4) were conducted with 1710 mm high and 110 mm thick masonry walls (Fig. 2). Walls 1–3 were 355 mm wide and Wall 4 was 230 mm wide. All four walls were reinforced on both faces with vertical NSM CFRP strips. The FRP strip was placed along the centerline of the flexural face, hence for Wall 4 (Fig. 2b), the strip alternately penetrated brick units and perpend joints in adjacent courses, while for the Walls 1–3 (Fig. 2a), the strip ran through the brick units. This allowed for investigating the influence of perpend joints and a different reinforcement ratio on the flexural behaviour of retrofitted masonry walls. It should be noted that unlike the monotonic pilot tests (Walls 1, 2 and 4), for the larger test specimens subjected to monotonic loading (Walls 5–10, 13 and 14) the FRP strip was placed only on one side (Fig. 3b) as it was thought that due to the small reinforcement ratios, not including FRP on the compressive face would

50 40

Applied Load (kN)

40

W-8

Applied Load (kN)

W-7

W- 14

30

W- 10

W-2

20

W-1

W-6

W-5

W-4

W- 12

30

20

10

10

W- 15

W- 11 W-3

W-9 0 0

20

40

60

Mid-height deflection (mm) Fig. 5. Load–displacement response for static tests.

80

0

0

20

40

60

Mid-height deflection (mm) Fig. 6. Load–displacement response for cyclic tests.

80

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result in a negligible difference in structural behaviour as well as saving money. However, as can be seen in Fig. 3c, the cyclically loaded specimens were reinforced on both sides (except for Wall 13). Wall 13 was unique in the sense that it was tested with reverse cyclic loading even though it only had NSM reinforcement on one face of the wall. Wall 13 was tested in order to show the wall’s behaviour in both its ‘strong’ and ‘weak’ directions.

2.4. Test setup

40

Applied Load (kN)

30

20

10

0 -80

-60

-40

-20

A vertical crack was observed on the compressive face (back face) directly behind the FRP strips on the tension face (front face) for Wall 10. This was thought to be due to the plane of weakness caused by the FRP strip on the opposite side. To avoid further complications due to the vertical crack interacting with the FRP strip, the spacing of the FRP on the back face was reduced by approximately 40 mm. Therefore, for Walls 11, 12 and 15 the centre-to-centre strip spacing was 535 mm on the front face and 456 mm on the back face (Fig. 3c).

0

20

40

60

80

-10

-20

Mid-height deflection (mm) Fig. 7. Load–displacement response for Wall 13.

40

The walls were simply-supported along their top and bottom edges with roller supports at the second courses from the top and bottom of the walls (refer Fig. 4). Walls 1–4 were subjected to four-point loading (Fig. 2) whereas Walls 5–15 were tested under three point loading (Fig. 3). For all the tests, the ends of the FRP strips were untrapped so that the final failure occurred once debonding propagated to the unloaded end. For the monotonic tests, the load was applied at their mid-height in one direction such that the FRP reinforcement was in tension. A stiffened roller that spanned the entire wall width was used to ensure that the load was applied uniformly across the entire width at the mid-height (refer Fig. 4). The monotonic load was increased slowly using a manually operated hydraulic jack until failure. While the walls were in effect loaded under ‘displacement control’, it was not possible to precisely control the loading rate since the hydraulic jack was operated by manual jacking. For reverse cyclic loading, the monotonic test arrangement was modified such that a reaction frame was constructed on both sides of the specimen and the hydraulic jack was attached to the centre course of bricks on each side, in order to push and pull the specimen back and forth. The cyclic tests were conducted by loading the walls in increments of 10–30% of the estimated ultimate deflection, as determined from the corresponding monotonic test which was always conducted beforehand. For each displacement increment two or three cycles of loading were applied with each cycle consisting of monotonically loading the wall in the positive direction until the target displacement was reached and then reducing the load to ‘‘zero’’ followed by monotonically loading the wall in the negative direction until the target displacement was reached and then reducing the load to ‘‘zero’’.

Wall 8 - 3 strips at 357mm

Applied Load (kN)

2.5. Test results

30

Wall 62 strips at 535mm

20

Wall 51 strip at 1070mm

10

0 0

20

40

60

80

Mid-height deflection (mm) Fig. 8. Load–displacement response showing effect of FRP strip spacing.

(a) Wall 5

Total applied load versus mid-height displacement plots for all walls are shown in Figs. 5–7. The wall bending tests results are summarised in Table 1, where Fmax refers to the maximum total applied load, Dult refers to the maximum displacement, emax refers to the maximum strain in the FRP strip during the test, Pexp refers to the maximum force in the strip and emax/erupt is the ratio, in %, of the maximum tensile strain recorded emax during the tests divided by the FRP rupture strain (Column 10 in Table 1). As can be seen from the result of Wall 13 (Column 9 in Table 1) which resisted 30.2 kN in its ‘strong’ direction whereas its ‘weak’ unreinforced direction strength was only 1.4 kN (Fig. 7), a significant increase in flexural strength over that of the corresponding unreinforced wall is possible even with the very small reinforcement ratio used in these walls. It should be noted that the flexural strength of all the walls was substantially higher than that required for normal seismic design situations and hence, highlights the effectiveness of the retrofitting scheme used for the bending tests. However, the increase in the strength due to retrofitting varied depending on the other test variables such as reinforcement ratio, applied axial load and strip spacing which will now be discussed. Due to the high reinforcement ratios in Walls 1–4 their failure mechanism was a combination of masonry crushing and IC debonding. In all other wall tests, failure

(b) Wall 6 Fig. 9. Failure mechanisms for Walls 5, 6 and 8.

(c) Wall 8

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pieces of brickwork attached. Another failure mechanism that becomes possible with large strip forces is tensile splitting of the masonry units along a vertical line on the compression face of the wall opposite to the NSM FRP strip. It appears that this can be avoided by keeping the compression face stresses well below the compressive design strength, fmc.

of the walls was by IC debonding. IC debonding describes the mechanism where the FRP debonds from the masonry starting at a flexural crack and then propagating away from the peak bending moment region where the first crack occurs towards the ‘unloaded’ end of the FRP strip (Willis et al. [3], Petersen et al. [4]). The majority of the FRP strips still had masonry attached to the FRP, indicating that the debonding failure surface is actually in the masonry material and not in the adhesive layer or at the FRP-adhesive interface. In some walls, after the diagonal herringbone cracks formed they propagated towards the perpend joints nearest to the FRP strip. Once the herringbone cracks reached the perpend joint the cracking followed the perpend joint to the next course of brickwork. Owing to the relatively short distance between bedjoints, as compared to the length required for full bond strength indicated by bond tests reported elsewhere (Kashyap et al. [18]), this herringbone cracking pattern continued to repeat itself from bedjoint to bedjoint as debonding progressed away from the point of maximum bending moment towards the top and bottom wall supports. At this point, the vertical in-plane shear strength of the masonry was unable to carry any further force so the strip and masonry commenced to fail with fairly large

2.6. Discussion of test results The following sections discuss trends observed in the experimental test results. Owing to the fact that repeat tests were not conducted, care must be taken when drawing conclusions from these tests. 2.6.1. Effect of FRP strip spacing Three walls (5, 6 and 8) were tested covering three different effective strip spacings (Figs. 8 and 9). The same total amount of FRP reinforcement (72 mm2) was used in each of the three test walls. From the test results (Table 1), it can be seen

(a) schematic of axial loading arrangement.

(c) view of axially loaded wall.

(b) load centering detail at top of wall.

Fig. 10. Axial loading details.

35 30

Wall 12 (σv = 0.1 MPa) 30

Wall 14 (σv = 0.2 MPa)

25

Applied Load (kN)

Applied Load (kN)

35

Wall 10 (σ σv = 0.1 MPa)

20 15

Wall 6 (σv = 0 MPa)

10

Wall 15 (σv = 0.2 MPa)

25 20 15 10

Wall 11 (σv = 0 MPa) 5

5

0

0 0

20

40

60

80

0

20

40

60

Mid-height deflection (mm)

Mid-height deflection (mm)

(a) Static tests

(b) Cyclic tests Fig. 11. Influence of pre-compression.

80

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M.C. Griffith et al. / Construction and Building Materials 49 (2013) 1032–1040

40

40

W-10 30

20

20

10

W-15 -80

-60

-40

-20

0

0

20

40

60

80

-10 -20

Applied Load (kN)

30

10

W-12

0 -80

-60

-40

-20

0

20

40

-10 -20 -30

-30 -40

-40

Mid-height deflection (mm)

Mid-height deflection (mm)

(a) σv = 0 MPa

(b) σv = 0.1 MPa 40

W-6 30

Applied Load (kN)

Applied Load (kN)

W-14

20 10

-80

-60

-40

-20

0

W-11 0

20

40

60

80

-10 -20 -30 -40

Mid-height deflection (mm)

(c) σv = 0.2 MPa Fig. 12. Effect of cyclic loading under different applied pre-compression.

Fig. 13. Failure mechanisms for static and cyclic tested Walls 10 and 12.

60

80

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M.C. Griffith et al. / Construction and Building Materials 49 (2013) 1032–1040

5

Wall 8, ρ = 6.1%

40

3 strip

Wall 7, ρ = 9.2%

pexp / Lper (kN/mm)

Applied Load (kN)

50

30

Wall 5, ρ = 6.1%

20

1 strip

4

Wall 8, ρ = 6.1%

3

Wall 5, ρ = 6.1% 2

Wall 9, ρ = 3.1%

1

10

Wall 9, ρ = 3.1%

Wall 7, ρ = 9.2% 0

0 0

20

40

60

80

100

120

0

20

40

60

80

Displacement (mm)

Displacement (mm)

(a) Load-displacement response

(b) FRP force normalized by Lper

Fig. 14. Influence of reinforcement ratio.

that there was a significant increase in both strength (approximately 114%) and displacement capacity (approximately 60%) when the FRP reinforcement was distributed more evenly. Wall 5 had a single strip which is unlikely to engage the entire width of the wall compared to Wall 6 and Wall 8 due to the smaller effective width, i.e. the width of the compressive zone which results in a total force equal to the tensile force in the FRP. Further understanding of this effect is given by considering the perimeter of the debonded failure plane, Lper, as defined by Kashyap et al. [18], where Lper = (b + 2 mm) + 2(d + 1 mm) and b and d are the thickness and width of the NSM FRP strip (in mm) and the ‘+1 and 2 mm’ terms account for the fact that the failure surface occurs in the brick unit and not at the FRP-adhesive interface. Here it is assumed that the failure surface is 1 mm away from the FRP strip faces. For example, the total surface area of FRP that was bonded to the masonry for Wall 8 (Lper = 3  [(4.8 + 2) + 2(5 + 1)] = 56.4 mm) was nearly double that for Wall 5 (Lper = 1  [(7.2 + 2) + 2(10 + 1)] = 31.2 mm). Finally, the most efficient use of the CFRP material was also achieved in Wall 8 where the maximum tensile stress recorded during the tests was observed to be 2055 MPa, or 89% of the rupture stress (Column 10, Table 1). 2.6.2. Effect of vertical pre-compression Six walls were tested with pre-compression ranging from zero to 0.2 MPa. The pre-compression was applied in force control to the top of the wall along the wall’s centreline as shown in Fig. 10 to maintain the effectiveness of the simple support at the wall’s top edge. Three walls were monotonically loaded (Walls 6, 10 and 14 – Fig. 11a) and three were subjected to cycli‘c loading (Walls 11, 12 and 15 – Fig. 11b), as shown in Table 1. As expected, increasing the axial load resulted in increased stiffness and maximum load but also with a decrease in the maximum displacement capacity. This is because the applied axial load makes the member stiffer as the flexural cracks are restrained by the combinations of the bond strength and the applied compressive load and hence, delaying the crack propagation. From Fig. 11 it can be seen that effects of pre-compression were more noticeable under monotonic loading than cyclic loading. 2.6.3. Effect of cyclic loading In Fig. 12, a mirror image of the monotonic load displacement curve has been superimposed on the cyclic test hysteresis plots in the negative direction for ease of comparison. As can be seen, the monotonic curves envelope the cyclic test results closely which suggests that cyclic load effects are not substantial. Overall, the cyclically loaded walls were marginally stiffer than the monotonically loaded walls; however, this was thought to be due to the placement of FRP on both faces of the cyclically loaded walls whereas the monotonic test walls only had FRP on one side. In comparison, cyclic loading resulted in a decrease in strength and displacement capacity by 6–8% and 15–20%, respectively. This was attributed to bond degradation due to cyclic loading (Fig. 13). 2.6.4. Effect of reinforcement ratio Fig. 14 shows that the amount of fibre reinforcement used affects the overall stiffness of a specimen and hence, the strength and displacement capacity. While evaluating the influence of reinforcement ratio, walls were selected such that other variables were constant. As can be seen in Fig. 14a, for the single strip configuration, doubling the reinforcement ratio led to an increase of 42% in strength whereas for the three strip configuration a 50% increase in reinforcement ratio resulted in a strength gain of only 12%. Further, it was noted that for the same reinforcement ratio (6.1% in Fig. 14a), the 3-strip configuration was roughly twice as strong and had 50% more displacement capacity than the corresponding 1-strip configuration. However, the relationship between strength and reinforcement ratio is clearly not linear.

To get a better understanding of the relationship between failure load and wall reinforcement patterns, consider Column 9 in Table 1 where it can be seen that the force in the FRP strip, Pexp, is smaller for walls with lower reinforcement ratio but when normalised by the bonded perimeter of NSM strip, Lper, the value of Pexp/Lper is reasonably consistent for all walls – ranging between 2.0 and 2.6 kN/mm, as shown in Fig. 14b. Thus, it can be deduced that the failure load is more closely related to the bonded surface area than reinforcement ratio. From Fig. 14a it can also be noted that the spacing of FRP strips has a greater influence on the overall behaviour of walls compared to reinforcement ratio and hence, is another key factor for efficient use of NSM FRP for retrofitting URM walls. This can also be observed from Column 10, Table 1, where emax/erup is not strongly dependent on reinforcement ratio whereas it is significantly affected by strip spacing.

3. Summary and closing remarks Fifteen NSM CFRP reinforced masonry walls were tested in vertical one-way bending to investigate the influence of strip spacing, reinforcement ratio, vertical pre-compression and cyclic loading on the flexural response of the walls. The test results confirmed that NSM reinforcement can be an efficient retrofit technique for increasing the vertical bending capacity of URM walls. In these tests they have been shown to provide increases in strength of up to 20 times the strength of the corresponding unreinforced wall. With respect to the test variables under investigation: Optimal spacing of FRP strips is beneficial not only in terms of strength and displacement capacity but can also help to avoid vertical in-plane shear failure or horizontal bending failure of the masonry between the FRP strips. For the same reinforcement ratio, multiple strips with smaller spacing produced stronger walls with larger displacement capacities (i) Optimal spacing of FRP strips is beneficial not only in terms of strength and displacement capacity but can also help to avoid vertical in-plane shear failure or horizontal bending failure of the masonry between the FRP strips. For the same reinforcement ratio, multiple strips with smaller spacing produced stronger walls with larger displacement capacities (ii) An increase in reinforcement ratio resulted in increased strength but also a corresponding reduction in displacement capacity. Closer inspection of the results revealed a nearly constant relationship between the failure load and the perimeter length of bonded FRP at a cross-section, Lper. This confirmed that for the same reinforcement ratio, more strength could be obtained by using smaller strips at closer spacing. The limit to this is that the FRP strips must have sufficient cross-sectional area to avoid tensile rupture failure. Furthermore, at the highest reinforcement ratio investigated in this study (14.2%), the wall’s failure mode appeared to be close to

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a vertical shear/splitting failure along the line of perpend joints parallel to the vertical FRP strip. Further study into the relation between the maximum bond force that the masonry can support along this potential perpend joint failure line may be warranted. Clearly, over reinforced walls can also be prone to compressive failure of the masonry. (iii) An applied axial load was seen to increase the flexural stiffness of walls but it had a minimal effect on the strength and displacement capacity of walls. (iv) Cyclic loading also was found to have a minimal effect on the strength of NSM CFRP reinforced masonry walls. Thus, for efficient use of FRP, it is recommended to choose optimal FRP strip spacing rather than increasing the reinforcement ratio as it provides better results for the overall behaviour of wall. Acknowledgements The research reported herein was supported by the Australian Research Council under grant number DP0879592 and was conducted at the University of Adelaide. The views expressed here are those of the authors and not necessarily those of the sponsor. The authors would also like to acknowledge the assistance to this project of the staff of the Chapman Laboratory numerous honours research students in the School of Civil, Environmental and Mining Engineering who assisted in carrying out the experimental work. References [1] Ingham JM, Griffith MC. Performance of unreinforced masonry buildings during the 2010 Darfield (Christchurch, NZ) earthquake. Aust J Struct Eng 2011;11(3):207–24.

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