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Proceedings of the Institution of Civil Engineers

Structures and Buildings 164 December 2011 Issue SB6

Pages 409–420 http://dx.doi.org/10.1680/stbu.2011.164.6.409

Paper 700028

Received 27/06/2007 Accepted 08/03/2011

Keywords: buildings, structures & design/composite structures/failures

ICE Publishing: All rights reserved

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

Structural response ofinterlocking compositemasonry slabj1 Waleed A. Thanoon PhD

Professor & Dean, College of Engineering and Architecture,University of Nizwa, Sultanate of Oman

j2 Yavuz Yardim PhDAssistant Professor, Engineering Faculty, Epoka University, Albania

j3 Mohd Saleh Jaafar PhDProfessor, Faculty of Engineering, University Putra Malaysia,Selangor, Malaysia

j4 Jamaloddin Noorzaei PhDProfessor, Faculty of Engineering, University Putra Malaysia,Malaysia

j1 j2 j3 j4

This study introduces a semi-fabricated composite floor slab system that consists of a precast inverted ribbed

ferrocement panel interlocked in situ with a brick–rib layer to form a composite slab. The effectiveness of the

interlocking mechanism of the composite slab was investigated under separate shear and flexural loadings. Ten

composite slab specimens having different shear connectivity between the two layers were cast and tested under

pure shear loading (push-off test). Six further slab specimens were cast and tested under two line loadings to explore

the structural response of the composite slab system under flexural loading. The flexural tests focused on the effect

of different brick layouts and orientation and thus different numbers of interlocking ribs in longitudinal and

transverse directions on the overall structural response of the composite slab. The results of push-off tests indicate

that the proposed interlocking system is as effective as using steel truss shear connectors in composite slabs. The

flexural test results in terms of load–deflection, crack pattern, ductility and failure loads indicate that the response of

the composite slab to flexural loading is satisfactory for use as a structural floor slab.

1. IntroductionReinforced concrete (RC) slabs are widely used in construction

because of their low cost, good performance and durability.

However, the cost of the formwork is high, ranging from 30 to

100% of the cost of the structure, depending on shape, size and

finishing requirements (Razali et al., 1993). In addition, the self-

weight of the concrete induces a high dead load on walls,

columns and beams.

Prefabricated floors are used in many parts of the world as an

alternative system to overcome formwork problems (cost and

delays in construction) and to control the quality of concrete.

However, prefabricated elements are very heavy and difficult to

transport and construct. In addition, concrete does not provide the

thermal insulation qualities desired for living accommodation.

Jointing connectivity is another problem observed in precast

construction, which leads to a less integrated structure.

To overcome these deficiencies, a large number of precast

systems have been developed. Pessiki et al. (1995) summarised

the use of 19 different precast structural floor systems that are

suitable for office building construction in different parts of the

world. Thin ferrocement panels have been used in floor construc-

tion for low-cost housing (Mansur and Ong, 1986; Omorodion-

Ikhimwin, 1983) due to their low cost and good structural

performance. The introduction of insulating sandwich panels has

increased the attractiveness of this type of construction. The

panels consist of thin layers of relatively higher strength material

sandwiching a thick core, normally of much weaker and lower

density material (Einea et al., 1994; Wright et al., 1987).

However, the high manufacturing and construction costs limit the

use of precast sandwich panels in construction. The use of

profiled steel sheeting as an integral part of RC deck slabs has

gained wide acceptance in many countries, especially those where

the cost of profiled steel sheeting is low (Kim and Youn, 2009;

Redzuan and Samuel, 2009). Such composite systems have been

used in bridges and long-span structures. The ultimate load

capacity of the profiled sheeting–concrete composite depends on

the shear bond characteristic and the slenderness parameters of

the profiled sheeting (Salmon et al., 1997). Cost and thermal

409

efficiency are some of the limitations of these systems. Profiled

sheeting–cement board composites, another recent development

in floor slab systems (Ahmed et al., 2002; Wan Badaruzzaman et

al., 2003), consist of profiled sheeting attached to a top layer of

dry board by simple mechanical connectors. Lightweight concrete

is used as an infill material to act as a sound insulator for the

floor. However, one of the limitations of this system is its low

stiffness, which results in large deflections and the development

of cracks in the finishing elements connected to the slab.

A semi-precast system known in Malaysia as a ‘half slab’ is

another development in floor slab construction (Yee, 2001a,

2001b). The technique employs a reinforced precast floor panel

that serves as permanent formwork for the composite with cast-

in-situ concrete. Steel lattice trusses projecting from the top of

the precast unit (Figure 1) are used to connect the two layers and

make the unit stiff during erection. The heavy weight of the full

slab, its low thermal efficiency and the additional cost of the steel

trusses needed to connect the two layers are some of the

disadvantages of the system.

Developing a new floor slab system to overcome the short-

comings of in situ concrete floor slabs and existing precast floor

systems is a challenging task for many researchers. Existing

systems have shortcomings such as long construction time, heavy

weight, dependency on heavy equipment at the job site, poor

thermal and sound insulation, high material wastage, dependency

on formwork, lack of structural integrity, jointing problems and

high cost.

This study introduces a semi-precast floor slab system, a

ferrocement–brick composite slab, to address some of the above

problems and to address the longitudinal shear failure commonly

observed in other composite slab systems (Thanoon et al.,

2010). The paper investigates the structural performance of the

ferrocement–brick composite slab under pure shear and flexural

short-term loadings. The main objective of the experimental

programme was to explore the efficiency of the proposed

interlocking system in connecting the two layers of the compo-

site. The deformational characteristics, ductility, ultimate

moment capacity and failure mechanism of the composite slab

are also presented and discussed.

2. Proposed ferrocement–brick compositesystem

The proposed floor slab is a semi-precast system that consists of

a precast inverted ribbed ferrocement panel (Figure 2(a)). This

layer provides a formwork for the bricks and cast-in-situ ribs.

Bricks are then laid in position (without the use of mortar) as

shown in Figure 2(b). The composite floor system is completed

by filling the in situ grooves between the bricks; these form the in

situ ribs as shown in Figure 2(c). Interlocking between the bricks,

the precast layer and the cast-in-situ ribs provides the necessary

shear transfer developed due to bending of the composite slab

(Figure 2(d)). The proposed interlocking concept does not require

any shear reinforcement (links or trusses) commonly used to

resist horizontal shear between two composite layers.

The precast ferrocement layer consists of wire mesh and steel

reinforcement to resist the flexural tensile stresses developed

during load transfer to the supporting beams or walls. The

thickness and reinforcement of this layer will mainly depend on

the span of the slab. The bricks and the in situ ribs provide the

necessary resistance to bear flexural compressive stresses.

One advantage of this system is that it is relatively light in weight

compared with an RC solid slab, which reduces the load

transferred to beams/walls. The masonry bricks act as a light

(especially voided brick), natural, cheap and effective insulation

material at the same time as partially resisting compression forces

developed due to bending of the composite.

On-site construction of the composite slab does not require any

heavy equipment to handle the ferrocement layer. Furthermore,

the construction does not require any formwork since the bottom

layer of the ferrocement is a precast unit that can be easily fixed

in position, using a simple crane, to provide a platform that acts

600 mm centre to centre

50–100 mm600/1200/1800/2400 mm

Figure 1. Half slab floor system

(a) Ferrocement precast layer

(b) Bricks laid in position

In situ ribs

(c) Ferrocement–brick composite floor

(d) Interlocked composite floor

Figure 2. Interlocking ferrocement–brick composite floor slab

410

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

as a formwork for the brick layer and the concrete ribs. The cold

joint problem usually observed in precast construction can also

be eliminated with this system. The floor slab units and support-

ing beams (or other slab units) will be integrated during casting

of the in situ concrete ribs.

3. Experimental programme

3.1 Test specimens

The effectiveness of the interlocking mechanism in transferring

shear stresses was initially investigated. Ten composite slab

specimens (750 mm 3 928 mm) with different shear connectiv-

ities between the two layers were cast and tested under pure shear

loading (push-off testing). The length of the specimens was

selected to account for brick–rib interaction. The specimens were

categorised into three types according to the way in which the

two layers were interconnected

(a) plain (P) specimens had no connectors between the two layers

(b) shear truss connectors were used to connect the two layers in

specimens of the second type (S specimens)

(c) the third type used an interlocking groove to connect the

layers (I specimens).

Specimens of the third type (interlocking system) were further

subdivided into groups IA, IB and IC to simulate different

interlocking depths by varying the depth of the precast and in situ

ribs. The depths of the interlocking grooves (precast ribs) were

fixed at 40, 30 and 20 mm respectively for specimens IA, IB and

IC (Figure 3).

The effectiveness of the proposed interlocking system in resisting

shear was explored by comparing the shear resistance of the

interlocked specimens with that of the specimens with steel truss

shear connectors. The shear connector used in S specimens was a

continuous steel truss placed at 250 mm centre to centre. The

truss members were made of 6 mm diameter mild steel bar of

depth 50 mm as shown in Figure 4.

The flexural response of the composite slab was investigated by

testing six slab specimens, each 1.5 m long, 0.75 m wide and

0.95 m thick. The specimens were tested under two-line loading.

The tests focused on the structural response of slabs with differ-

ent interlocking ribs as a result of different brick layouts. The

specimens were arranged in three groups (A, B and C), with each

group having two identical specimens. The brick arrangements of

each group are shown in Figure 5. In all the specimens, the depth

of the ferrocement layer (including rib depth) was 60 mm; this

was reinforced with two layers of 1.2 mm diameter wire mesh of

12.7 mm 3 12.7 mm opening and 4T10 mm diameter steel rein-

forcement. The hollow brick layer was 65 mm thick (brick size

216 3 100 3 65 mm). The surface areas and volumes of the

70 215 70 215 70 215 70

50

50

50

AA

100

750

Plan

In situ ribs

Bricks Bricks Bricks

Ferrocement precast layer

110 mm

45 mm

Variable 25, 35, 45 mm

Interlocking depth,variable 40, 30, 20 mm

Section A–A for interlocking group

Figure 3. Plan and cross-section details for tested specimens

(dimensions in millimetres)

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Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

bricks with respect to total surface area and total volume of the

specimen are presented in Table 1.

In group A, the surface area of the bricks was equal to 58% of

the total surface area. The brick layout in this group resulted in

three longitudinal ribs (total width 150 mm) and six ribs of

70 mm width in the transverse direction. Group B specimens

were identical to group A specimens except that the layout of the

bricks produced four ribs in the longitudinal direction (with the

same total width of 150 mm as group A). The surface area of the

bricks in group C was equal to 70% of the total surface area. The

brick layout was arranged to have four ribs 25 mm thick (total

width 100 mm) in the longitudinal dimension and five ribs of

60 mm thickness in the transverse direction. A sample of the

analytical calculations carried out to estimate the flexural capa-

cities of these specimens are presented in the appendix.

3.2 Preparation of the specimens

For both the brick and ferrocement layers, ordinary Portland

cement and natural sand were used in a ratio of 1:3 with a water/

cement ratio of 0.5. The 28-day average cube strength of this mix

was 30 MPa. The tensile strength of the wire mesh and steel

reinforcement (tested using a universal test machine) was found

Steel trusses

AA

250

250

250

Plan

50

58

110

Ferrocement precast layer

Section A–A

Figure 4. Details of slabs with shear truss connectors

(S specimens) (dimensions in millimetres)

Section A–AGroup B

50200

25

1500A

Brick

750

70 70A

216

953530

Steel bar Wire mesh

Section A–AGroup A

50

300

50

1500A

Brick

750

70 70A

216

953530

Steel bar Wire mesh

Section A–AGroup C

25216

25

1500A

Brick

750

60 60A

300

953530

Steel bar Wire mesh

Figure 5. Brick layouts for flexural testing (dimensions in

millimetres)

Group Surface area

of bricks: %

Brick volume:

%

No. of ribs in

longitudinal

direction

No. of ribs in

transverse

direction

Rib width in

transverse

direction: mm

Rib width in longitudinal

direction: mm

Edge Interim

A 58 40 3 6 70 50 50

B 58 40 4 6 70 50 25

C 70 48 4 5 60 25 25

Table 1. Details of flexure test specimens

412

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

to be 300 MPa and 415 MPa respectively. Initially, the ferroce-

ment layer was cast after preparing the wire mesh and steel

reinforcement as shown in Figure 6(b). Polystyrene was used to

obtain the profile of the precast ferrocement layer. The ferroce-

ment layer was cured for 7 days and then the polystyrene pieces

were removed and replaced with bricks. The grooves between the

bricks (the in situ ribs) were filled with mortar to complete the

casting process (Figure 6(d)).

4. Test set-up

4.1 Push-off test

All ten specimens were tested under pure shear load. A special

frame was fabricated for the test. The frame was fixed to a strong

floor as shown in Figure 7. A steel beam (I section) with thick

steel plates was used to apply load using a hydraulic jack system.

The jack itself was fixed to the floor using steel plates and high-

strength bolts. The precast layer at the bottom was restrained

from movement in the direction of the applied load using steel

plates connected to the strong floor with high-strength bolts. A

horizontal distributed load was applied incrementally until failure

of the specimen. As shown in Figure 7, a roller support was

placed on the top of the specimen to provide rotational equili-

brium due to the eccentricity of the loading and to prevent

instabilities inherent with this type of test (Gohnert, 2003). Three

dial gauges were fixed to measure the slip of the topping layer

relative to the fixed precast layer.

4.2 Flexure test

The six specimens were tested as simply supported slabs over a

1.5 m span with a concentrated two-line load applied as shown in

Figure 8. This arrangement was used to obtain a shear span ratio

greater than 5.0 (av/d ¼ 5.8, where av is shear span and d is

effective depth) to ensure flexural failure before shear failure.

The load was applied gradually using a hydraulic jack. At every

increment of the load, readings of the dial gauges and strain

gauges were recorded until failure of the slab. Deflection in the

middle region of the slab was continuously monitored using dial

gauges and linear variable differential transducers (LDVTs). The

locations of the cracks were marked with progression of the

applied load.

(a) Rib casting (b) Casting ferrocement layer

(c) Removing polystyrene layers (d) Placing brick and casting the in situ ribs

Figure 6. Casting specimens

413

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

5. Results and discussion

5.1 Shear test

Figure 9 shows the relation between applied shear load and slip

between the two layers of the composite slab. A very small slip

(1–2 mm) was observed in all the tested specimens apart from

those with steel truss shear connectors (S specimens). In general,

the load–slip characteristics of S specimens and I specimens

were similar, and also quite similar to those reported by Moy and

Tayler (1996) and Gohnert (2000). Before failure there was a

sudden increase in the slip followed by a brittle type of failure.

The maximum slip and failure loads of all the specimens are

listed in Table 2. The efficiency factors – defined as the ratio of

experimental ultimate shear load of specimens with shear con-

nectors to that of specimens without any shear connectors – were

calculated and are also shown in the Table 2. The efficiency

factor of specimens with dual steel truss shear connectors was 3.1

and the specimens with the proposed interlocking system showed

comparable results with efficiency factors of 2.5–2.9 (average

2.75). This indicates that the proposed interlocking mechanism is

as effective as steel trusses in resisting shear stresses and could

thus be used to replace steel trusses with consequent reductions

in cost.

For specimens without a shear connector, failure was sudden and

without any signs of cracking (Figure 10a). The in situ topping

layer slid on the precast layer when the shear stress equalled

0.17 MPa. Specimens with dual shear truss connectors showed a

longitudinal crack along the steel truss (Figure 10(b)) at a load

equal to 90% of the failure load. The crack width increased with

increases in applied load until the mortar crushed around the

trusses without failure of the steels.

The interlocking group of specimens failed by shearing off of the

interlocking ribs as shown in Figures 10(c) and 10(d). However,

one specimen suffered a bearing failure in the rib at the vicinity

of the applied load (Figure 10(e)). All the specimens showed

brittle type of failure. However, the specimens with steel truss

shear connectors showed much higher slips and cracks before

their sudden failure.

Load

Interlock depth

Reaction

Fix

Figure 7. Pure shear test set-up

Hydraulic jack

I beam

545 mm 310 mm 545 mmL 1500 mm�

Two-line load test (flexural test)

Figure 8. Flexure test set-up

414

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

5.2 Flexure response

The load–deflection characteristics for the tested slab specimens

are shown in Figure 11. All the slabs behaved in an elastic

manner before cracking, after which the stiffness of the speci-

mens was reduced as indicated by the load–deflection curves.

These responses are similar to the well-known RC slab behaviour.

Table 3 summarises the cracking loads of the ferrocement, loads

at which main steel reinforcement yielded (according to measured

yielding strain of steel reinforcement), ultimate loads at failure

and ductility (defined here as the ratio of deflection at ultimate

load to the deflection at yielding load) of each specimen.

All the specimens show high ductility (4.2–6.5), which provides

sufficient warning before failure. The high ductility of the speci-

mens is due to the presence of the ferrocement layer and the

uniform transfer of stresses through the interlocking mechanism

provided between the two layers of the slab.

The load–deflection curves of all the tested specimens are

combined in Figure 12 for comparison. Different brick layouts

and orientation resulted in different flexural responses. Although

group A and B specimens had similar brick surface areas,

transverse ribs and total width of longitudinal ribs, group B

specimens exhibited stiffer responses than group A specimens.

However, the ultimate failure load for both groups was almost

equal (only 7% difference). Moreover, the ductility of the group

B specimens was 20% higher than that of group A specimens.

These responses might be due to the number of longitudinal ribs,

which probably enhances the flexural response, but more tests are

needed before more solid conclusions can be drawn.

The load–deflection response of the group C specimens showed

the lowest ultimate capacity. This is due to the completely

different brick layout and orientation in this group – the total

width of the longitudinal ribs was 100 mm, 33.3% less than that

of groups A and B. The layout and orientation of the bricks in

390

360

330

300

270

240

210

180

150

120

90

60

30

00 1 2 3 4

Deflection: mm

Load

: kN

P

S

IC

IB

IA

Figure 9. Shear test: load–slip characteristics of tested specimens

Specimen Connector type Load: kN Maximum

slip: mm

Efficiency

factor

P1 None 114.8 1.2 1.0

P2 None 107.6

S1 Dual steel trusses 358.8 3.6 3.1

S2 Dual steel trusses 342.9

IA1 40 mm interlocking groove 300.5 2.1 2.5

IA2 40 mm interlocking groove 262.9

IB1 30 mm interlocking groove 290.0 2.2 2.6

IB2 30 mm interlocking groove 293.0

IC1 20 mm interlocking groove 339.0 2.6 2.9

IC2 20 mm interlocking groove 314.8

Table 2. Details of push-off test results at failure

415

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

this group reduced both the number and size of the transverse

ribs. The ultimate capacities of group A and B specimens were

respectively 21% and 29% higher than those of group C.

However, the surface area of the bricks in group C was 20%

higher than the other two groups; this leads to a reduction in

weight of the panel and better thermal properties.

All specimens showed a similar cracking load of 12 kN (about

30% of the ultimate load). The cracks started in the ferrocement

layers and gradually extended upwards with increases in applied

load. The crack patterns of different specimens at failure are

shown in Figure 13. The majority of cracks are concentrated at

the peak moment region. No longitudinal cracks were observed

between the two layers of the slab throughout the entire loading

process, indicating that the slab specimens behaved in a fully

composite manner until failure. Similar failure patterns were

found in all the slab specimens: widening of flexural cracks that,

at the moment of failure, extended upwards before crushing the

top part of the slab at this region as shown in Figure 13. The

widest crack was observed in the group C specimens. The failure

patterns are quite similar to those observed in RC one-way slabs.

The ultimate moments recorded for all specimens are also

presented in Table 3. The values range from 12 to 15 kNm per

metre width of slab, which is considered adequate to resist an

equivalent characteristic uniformly distributed imposed load of

3–4 kN/m2:

The analytical calculations presented in the appendix (BSI, 1997;

Chakrabarti et al., 1988) show that the estimated failure loads for

different slab specimens were 40.7 kN for groups A and B and

40.3 kN for group C. These capacities were calculated without

using material safety factors. The calculations give good esti-

mates for the ultimate load compared with the average ultimate

experimental load found in group A and B specimens. However,

for group C specimens, the analytical calculation gives an

ultimate load 20% higher than that found experimentally. A more

comprehensive method is required to estimate the analytical

capacity of the composite slab.

6. ConclusionThis paper has introduced a semi-fabricated composite system for

a floor slab. It consists of a precast inverted ribbed ferrocement

(a) Type P specimen

(c) Type IC specimen (e) Type IA specimen(d) Type IB specimen

(b) Type S specimen

Figure 10. Failure mechanisms of different specimens in push-off

shear tests

416

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

panel interlocked in situ with bricks and a mortar layer to form a

composite unit. The interlocking system was proposed to address

the longitudinal shear failure commonly observed in other com-

posite slab systems.

The results of push-off tests indicate that the proposed interlock-

ing system is as effective as using steel truss shear connectors in

a composite slab. The flexural test results in terms of load–

deflection, crack pattern, ductility and failure loads indicate that

the response of the composite slab to flexural loading is

satisfactory for use as a structural floor slab. The slab can resist a

bending moment of 12–15 kNm/m, which is considered adequate

to resist a characteristic imposed load of 3–4 kN/m2: All the

specimens showed high ductility in the range 4.2–6.5, providing

sufficient warning before failure. The high ductility is attributed

to the ferrocement layer and the interlocking mechanism provided

between the two layers of the slab.

The failure patterns are quite similar to those observed in RC

one-way slabs. Although the specimens with the proposed inter-

locking system failed in a brittle manner under pure shear load,

the flexural test results did not show any sign of longitudinal

cracking between the layers of the composite over the entire load

range, indicating that the slab behaved in a fully composite

manner until failure. The brittleness of the shear failure does not

affect the flexural performance of the composite due to the

ductile failure of the composite slab, as reflected in the high

ductility ratios observed in the test results. More tests are

required (full-scale tests, jointing systems, etc.) before the

proposed system may be adopted in construction.

7. Appendix

7.1 Moment capacity

The ultimate moment capacity of the ferrocement–brick compo-

site slab may be estimated using equilibrium equations and

compatibility of strains in the composite layers. BS 8110 (BSI,

1997) and the work of Chakrabarti et al. (1988) were used to

calculate the ultimate moment capacities of the composite slabs

in different groups. It was assumed that

(a) the slab is fully composite �brick ¼ �rib (where � is

deformation)

(b) plane sections remain plane after bending

45

40

35

30

25

20

15

10

5

0

Load

: kN

0 2 4 6 8 10 12 14 16 18 20Mid-span deflection: mm

(a)

Specimen A1Specimen A2

Yielding A1Yielding A2

Cracking

45

40

35

30

25

20

15

10

5

0

Load

: kN

0 2 4 6 8 10 12 14 16 18 20Mid-span deflection: mm

(b)

Specimen A1Specimen A2

Yielding B1

Yielding B2

Cracking

45

40

35

30

25

20

15

10

5

0

Load

: kN

0 2 4 6 8 10 12 14 16 18 20Mid-span deflection: mm

(c)

Specimen A1Specimen A2

Yielding C1

Yielding C2

Cracking

Figure 11. Load–displacement characteristics: (a) group A

specimens; (b) group B specimens; (c) group C specimens

50

45

40

35

30

25

20

15

10

5

0

Load

: kN

0 2 4 6 8 10 12 14 16 18 20Mid-span deflection: mm

Specimen A1 Specimen A2Specimen B1 Specimen B2Specimen C1 Specimen C2

Figure 12. Comparison of the deformation response of different

specimens

417

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

Specimen Cracking

load: kN

Yielding

load: kN

Ultimate

load: kN

Ultimate

moment: kNm

Ductility

�u/�y

A1 12.5 27.4 41.2 11.3 4.75

A2 13.0 24.0 39.3 10.8 5.60

B1 12.5 28.8 43.1 11.75 5.65

B2 12.0 25.0 43.1 11.75 6.70

C1 12.3 19.2 33.5 9.13 5.10

C2 11.5 25.0 33.4 9.1 4.20

Table 3. Details of flexure test results

CL

CL

CL

Crushing

(a)

(b)

(c)

Figure 13. Failure mechanisms of (a) group A, (b) group B and

(c) group C specimens

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Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

(c) the effect of the transverse ribs in stiffening the longitudinal

ribs can be ignored.

The material characteristics are given in Table 4. The estimated

ultimate load was determined as follows. Referring to Figure 14,

the equilibrium equation C ¼ T is used to define the depth of the

neutral axis x (where C is compressive force and T is tensile

force).

0:67 fbu(0:9x)bbrick þ 0:67 f cu(0:9x)brib

¼ 0:91(As f y þ Aw fyw)

¼ [0:673 103 6003 (0:9x)]

þ [0:673 303 1503 (0:9x)]

¼ 0:91[(3143 415)þ (1813 300)]

where fbu is brick compressive strength, fcu is mortar compressive

strength, As is area of steel reinformcement, Aw is area of wire

mesh and fyw is tensile strenght of wire mesh, and bbrick and brib

are width of bricks and ribs respectively.

3618xþ 2713:5x ¼ 167 995:1

\x ¼ 26:5 mm

Mu ¼ 0:91(As f y) d1 �0:9x

2

� �

þ 0:91(Aw f yw) d2 �0:9x

2

� �

Mu ¼ 0:91(3143 415) 78� 0:93 26:5

2

� �

þ 0:91(1813 300) 78� 0:93 26:5

2

� �

Mu ¼ 11:1 kNm (for specimens in groups A and B)

where Mu is moment capacity and d1, d2 are effective depth with

respect to steel and wire mesh, respectively.

The total failure load ¼ 40.7 kN

The axial compression force in the bricks and ribs are calculated

as

Cbrick ¼ 95:8 kN

Crib ¼ 57:6 kN

where Cbrick and Crib are the compressive force in the bricks and

ribs, respectively.

7.2 Interlocking stresses

The total horizontal shear force between the two layers at ultimate

moment of the panel is equal to the total tension or compression

force in the layers (maximum at mid-span) (Salmon et al., 1997).

So the maximum compression force C at mid-span is 153.4 kN.

This force will be equal to zero at support as M ¼ 0. The average

compression force along half of the span can be assumed to be

Vava ¼ C=2

Then the stresses in bricks and ribs may be estimated as follows.

Shearing area of bricks ¼ 2163 2.53 600 ¼ 324 000 mm2

Shearing area of ribs ¼ 3(750 3 70) + (3 3 216 3 50 3 2.5)

¼ 238 500 mm2

Mortar compressive strength fcu: MPa 30

Steel yielding strength fys: MPa 415

Tensile strength of wire mesh fyw: MPa 300

Area of steel reinforcement As: mm2 314

Area of wire mesh Aw: mm2 181

Brick compressive strength fbu: MPa 10

Table 4. Material characteristics

Brick

Wire meshAs2314 mm�

x C

T

35 mm

60 mmN.A

z

50 mm3 100 mm

300 mm�

�50 mmIn situ rib

Figure 14. Cross section of slab specimens in group A

419

Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

The shear stress in bricks is

vuðbrickÞ ¼95:83 0:53 103

324 000¼ 0:15 MPa

and shear stress in the ribs is

vuðribÞ ¼57:63 0:53 103

238 500¼ 0:12 MPa

The maximum bearing stress at the interlocking groove (having

30 mm depth) may be calculated as

95:83 103

6003 30¼ 5:5 MPa

which is less than the crushing capacity of the brick used

(10 MPa).

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Structures and BuildingsVolume 164 Issue SB6

Structural response of interlockingcomposite masonry slabThanoon, Yardim, Jaafar and Noorzaei

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