Load Capacity of Dry-Stack Masonry Walls

Load Capacity of Dry-Stack Masonry Walls
H. C. Uzoegbo1, R. Senthivel2 and J. V. Ngowi3
The construction industry is acknowledging the strong
need to accelerate the masonry construction process, as the
traditional method is labour intensive, and hence slower,
due to the presence of a large number of mortar joints
[Anand and Ramamurthy (2000)]. Early attempts were
made to increase the size of masonry units (block instead
of brick), thereby reducing the number of mortar joints,
wherein the use of bedding mortar imposed constraints on
the number of layers to be constructed in a day. The need
for further acceleration of the rate of construction led to
the elimination of bedding mortar and the development
of non-conventional methods of masonry construction
techniques such as the Hydraform Dry-Stack Block masonry [Agre´ment South Africa (1996), Uzoegbo, (2001),
Uzoegbo and Ngowi, (2003) and Uzoegbo, (2003)]. The
technique of dry stacking masonry units in construction
has existed in Africa for thousands of years. The Egyptian
pyramids and the Zimbabwe ruins, a capital of ancient
Shona Kingdom around 400AD, are good examples. Innovative mortarless system has improved with time since
mid 1980’s and is now more competitive in the market
than before. Several institutions in America, Africa and
Asia are now involved in the development of this technology. However, little attention has been given to research
on the structural behavior of the systems. Hydraform
of South Africa, Azar and Spurlock both of Canada and
Haener of USA are among many companies that are
currently developing and marketing dry-stack masonry.
Dry-stack masonry units of different geometry, sizes
and interlocking features have been developed in recent
years. In conventional masonry, mortar is used for bonding the masonry units. Dry stacking relies mainly on the
mechanical interlocking features of the units, which assist
alignment and provide stability during construction. Dry
stacking reduces the requirement for skilled labour and a
costly bonding material like cement and allows floor and
roof Loadings to be applied immediately upon completion of walls. It reduces building costs due to savings in
construction time. Overall savings of up to 27% compared
to conventional masonry have been reported [Agre´ment
South Africa (1996)]. The savings are mainly due to savings in cost of mortar, the block units and construction
Professor, School of Civil & Envir. Engg., University
of the Witwatersrand, Johannesburg, South Africa.
2
TMS Member, Post-Doctoral Research Fellow, School
of Civil & Envir. Engg., University of the Witwatersrand, South Africa.
3
Ph.D. Student, School of Civil & Envir. Engg., University of the Witwatersrand, Johannesburg, South
Africa.
time [Uzoegbo (2001)]. The behavior of interlocking
block walls have been studied and reported by several
researchers [Drysdale and Gazzola (1991), Gazzola and
Drysdale (1989), Harris and Hamid (1992), Harris and
Hamid (1993), Hatzinikolas, Elwi and Lee (1986), Harris and Hamid (1993), Oh (1994)]. The University of the
Witwatersrand in collaboration with Hydraform Africa
Ltd is currently investigating the structural behavior of a
dry-stack masonry developed by Hydraform [Agre´ment
South Africa (1996)]. This paper presents an investigation of full-scale dry-stack interlocking masonry walling
system and wallettes constructed in the laboratory using
Hydraform interlocking blocksTM and tested under axial
compression, lateral tension and flexural bending loads.
The interlocking block dimensions are 220 mm (8.66 in.)
width, 115 mm (4.53 in.) height and 220 mm (8.66 in.)
length (Figure 1). The units are designed to fit into the
groove provided by the unit under it. The units were made
in compressive strength of 5 MPa (725.2 psi), 7 MPa
(1,015.3 psi), 9 MPa (1,305.34 psi), 12 MPa (1,740.5 psi)
and 23 MPa (3,335.87 psi). The test specimens were constructed in the laboratory using the construction method
described in the Hydraform block masonry construction
manual [Agre´ment South Africa (1996)].
EXPERIMENTAL PROGRAM
Axial Compression
Wall panels were built using Hydrafrom interlocking blocks of the following block unit strengths: 5 MPa
(725.2 psi), 7 MPa (1,015.3 psi), 9 MPa (1,305.34 psi),
12 MPa (1,740.5 psi) and 23 MPa (3,335.87 psi). The
test specimens were constructed for each block grade
and tested in the Macklow–Smith machine platen that is
mounted on a hydraulic Ram. The size of each wall panel
1
TMS Journal September 2007
Figure 1—Hydraform Interlocking Dry-Stack BlockTM
(1 mm = 0.0394 in.)
41
Figure 2—Details of Test Panel and Dial Gage Positions (Number Represents Dial Gage Positions (1 m = 3.28 ft)
a) Front View
b) Side View 1
c) Side View 2
d) Details of Upper Part of Side View 2
Figure 3—Modes of Failure (1mm = 0.0394 in.)
42
TMS Journal September 2007
Table 1. Compressive Strength of Wall Panels (1 kN = 101.97 kp, 1 mm = 0.039 in.)
Ultimate Compressive load
Maximum lateral Displacement
Type of Wall panel
(kN)
(mm)
5MPa Units, dry-stack
595
2.3
9MPa Units, dry-stack
721
10.0
12MPa Units, dry-stack
938
3.4
12MPa Units, bonded
1,553
4.0
23MPa Units, dry-stack
1,360
40.0
was 3 m (118.11 in.) long, 2.5 m (98.43 in.) high and 220 mm
(8.66 in.) thick. The starter course was laid in mortar (1:6)
and cured for three days without load. The rest of the courses
were dry-stack, with the units at the edges of the wall laid
in EverbondTM (an adhesive type bonding agent). The last
three top courses were also laid in EverbondTM. The end
vertical strips were also bonded as indicated by the shaded
area in Figure 2. The top surface of the wall was capped with
mortar and checked by spirit level. The structures were tested
at 14 days. A 3 m (118.11 in.) span steel beam was used to
spread the load evenly at the top of the wall. Dial gauges
were placed in positions indicated as shown in Figure 2.
Axial compression load was applied at a rate of 2 kN/min
(203.94 kp/min). At each interval the lateral displacement
was recorded from the dial gauges and the corresponding
load from machine control panel was read.
Wall panel tests indicate that the onset of failure is
characterized by the formation of a vertical crack (less
than 3 mm (0.11 in.)) parallel to the axis of loading along
the mid section of the wall (Figure 3a). At ultimate state,
cracks also appeared on the faces and edges of the specimen
(Figure 3b). The appearance of the cracks was accompanied in each test by a loud snap and sudden reduction in the
magnitude of the load applied, which quickly reduced to
zero. The main failure plane was along the vertical joint of
the interlocking mechanism as shown in Figures 3b and 3c.
However, failure of samples made with low-strength units
(less than 5 MPa (725.2 psi)) walls was characterized by
a local crushing of the top courses as shown in Figure 3c.
The average ultimate load at the point of failure and corresponding maximum average lateral displacement (normal
to the axial load) for each wall type is shown in Table 1 and
also shown in non-dimensional form in Figure 4 and 5 respectively, where the load is normalized with respect to the
maximum failure load of the five walls with different unit
strengths and the displacement is normalized with respect
to the displacement of the maximum failure load.
A control wall panel using conventional blocks was
built with 12 MPa (1,740.5 psi) units and class II (1:6)
mortar. The structure was tested at 28 days. The mortar
cube strength was 5.6 MPa (812.21 psi) at 28 days. Figure 6
shows the load capacity of both the conventional wall (units
bonded with mortar) and the corresponding wall with a
dry-stack system using 12 MPa (1,740.5 psi) units. The
comparison is shown in non-dimensional form where the
load is normalized with respect to the failure load of the
conventional wall.
Figure 4—Normalized Failure Loads of Dry-Stack Walls (1 MPa = 145.038 psi)
TMS Journal September 2007
43
Figure 5—Normalized Maximum Lateral (Normal to Load) Displacements of Walls
Figure 6—Normalized Failure Loads of Dry-Stack and Conventional Walls
The test results were used to establish a relationship
between the unit strength and the masonry panel strength.
The average compressive strength of the dry-stack wall
panel fpanel as a function of the masonry unit strength, fcu
is given as:
f panel = φm 0.15 f cu + 1
where
fpanel =
φ =
fcu =
masonry panel strength, MPa (psi)
safety factor on material, 0.9
masonry unit strength, MPa (psi)
(1)
The above expression was compared with test results
as shown in Figure 7. Comparison between the experimental and computed results shows good agreement. The
proposed expression is valid for panels consisting of blocks
44
with compressive strength in the range of 5 - 25 (725.2
– 3625.9) MPa (psi).
Lateral Loading Test
Lateral loading tests were conducted on a full-scale
single roomed structure constructed in the laboratory
(Figure 8). Three specimens with different span lengths
were tested: short span (Ss = 1.5 m (59.1 in.)), medium
span (Ms = 3 m (118.11 in)) and long span (Ls=6 m (236.22
in.)). All the specimens were 2.45 m (96.46 in.) height
and constructed using Hydraform interlocking blocks of 9
N/mm² (1,305.34 psi) units strength. The length: thickness
(L/t) ratio ranged from 6.8 to 27 and the h/t ratio was 11 for
all the tests. In order to provide level surface to keep the
wall aligned vertically and horizontally the starter course
was laid in class II cement-sand (1:6) mortar and cured
TMS Journal September 2007
Figure 7—Strength of wall Vs. Block Unit (1 kN = 101.97 kp 1 MPa = 145.038 psi)
Figure 8—Full-Scale Single Roomed Test Specimen (3 m span (1 m = 3.28 ft))
for 3 days without load. Courses in the middle area of the
panels were dry-stack. The top three lintel courses were
also laid in mortar to provide a ring beam to resist uplift.
The structure was cured for 14 days before testing.
The experimental test set-up consisted of a loading
frame bolted to the laboratory floor as a support for application of lateral load to the wall as shown in Figure 9.
The load was applied by means of a reinforced water bag
of dimensions 1 m × 1 m (39.37 in. × 39.37 in.), which
is inflated by means of water pressure from the domestic
water reticulation system. A pressure of 1 kN/m² (0.15 psi)
over the 1m² (1,550 in.2) area of the power bag therefore
TMS Journal September 2007
translates to a force of 1 kN (101.97 kp) in the wall. The
power bag is connected to a pressure transducer and dial
gauge for pressure readings. A one-way valve is also
connected to the system to eliminate back pressure. The
built test structure is shown in Figure 8. The arrangement
of reinforced water bag and position of strain gauges are
shown in Figure 9.
The modes of failure are shown in Figure 10 (a – d)
for all three spans. The failures of the walls were mainly
due to excessive deflection at middle of wall (where load
was applied through water power bag on 1 m2 (1,550 in.2)
area) and mid span of the upper half of the wall. The failure
45
Figure 9—Experimental Set-Up For Lateral Load Test and Position of Dial Gauges
a) Short Span
b) Medium Span
c) Long Span (Side View)
d) Long Span (Front View)
Figure 10—Modes of Failure
includes opening/splitting of the head joints (at upper half of
the walls) and inclined shear cracks along the head and bed
joints (at lower half of the walls). Failure of the interlocking
mechanism by shear of some units was also observed in the
area of maximum deflection (i.e. the area at the upper half of
the wall). The overall failure mode was similar to the yield
line pattern in laterally loaded reinforced concrete slab. It has
been observed that the starter course tends to lift 5 to 12 mm
(0.197 to 0.472 in.) before failure of the specimens. A gradual
movement of the wall in response to load increase character46
ized the general load deflection behaviour of the wall panels.
The ultimate load at failure was 7.67 kPa (1.1 psi), 3.6 kPa
(0.52 psi) and 1.17 kPa (0.17 psi) for the short, medium
and long spans respectively. The corresponding mid span
deflection was 84.1 mm (3.31 in.), 86.7 mm (3.41 in.) and
63.6 mm (2.5 in.). The load-deflection curves of all three
spans are shown in non-dimensional form in Figure 11 where
the loads are normalized with respect to the peak load of the
short span and the deflections are normalized with respect
to the short span deflection at the failure load.
TMS Journal September 2007
Figure 11—Normalized load-deflection for three different spans (1 m = 3.28 ft)
Figure 12—Lateral Load resistance capacity of dry-stack masonry (1 kPa - 0.145 psi, 1 m = 3.28 ft)
The results from the lateral load tests were used to
establish the relationship between load carrying capacity
and the length of the panels in Figure 12. The test results
suggest that the lateral load capacity of a dry-stack wall
panel is dependent on the length of the wall panel (longitudinal direction), type, geometry and characteristics of
interlocking mechanism.
and dry stack joints were tested. The wallets of each load
case were tested under four point loading and the load was
applied by using a hydraulic jack. A load cell to measure
the applied load and a dial gauge was used to monitor the
deflection at the middle of the specimen.
Flexural Strength Tests
The dimensions of the wallettes were 1,495 mm
(58.86 in.) high (13 courses) and 960 mm (37.8 in.) long
(4 blocks). Three specimens were tested under four-point
load parallel to the bed joints. Figure 13 shows the test
specimen and experimental set-up. To minimize the friction
at the bottom of the specimens, pieces of maisonette with
rollers between them were first laid and then wallettes built
on top. No special treatment such as curing was required.
The wallettes were dry stacked and tested immediately.
The specimens were loaded under four point loading and
the lateral pressure was applied by means of a hydraulic
jack supported by a rig frame. Rubber pad (packing) was
Six scaled models (wallettes or prims) were constructed using Hydraform interlocking blocksTM of 9 MPa
(1,305.34 psi) unit strength, which is similar to the one used
in the lateral load tests. Attempt was made to establish the
flexural strength of the dry-stack masonry with bending
parallel (vertical bending) and perpendicular (horizontal
bending) to the bed joints respectively. Three specimens
were tested for each load case. The size of the specimens
was chosen such a way that adequate number of courses
TMS Journal September 2007
Load (Bending) Parallel to the Bed Joints
47
Figure 13—Experimental Set-up for Flexure Test (Parallel to the Bed Joints) (1 mm = 0.039 in.)
Figure 14—Failure Mode (Parallel to the Bed Joints)
introduced between the bearing and the specimen. The
deflection of the specimen at the middle was monitored
using a dial gauge.
Figure 14 shows the typical modes of failure of the
specimens, which was characterised by the gradual deflection with load increase. This was also accompanied by the
gradual opening of the bed joint above the mid-section
(between the loading points) stretching across the entire
48
length of the specimen. An average ultimate load of about
1.5 kN (152.96 kp) was attained at the point of failure. At
the failure load, the maximum opening of the bed joints
due to the rotation of the units about vertical axis was about
50 mm (2 in.) resulting in the sliding of the units out of the
interlocking mechanism. The specimen continued to deflect
with the decrease of the lateral pressure resistance with further rotation of the units before the loading was stopped. All
the head joints practically remained closed. Failure of the
TMS Journal September 2007
a) Side Elevation
b) Side View
Figure 15—Experimental Set-up for Flexure Test (Perpendicular to the Bed Joints)( 1 mm = 0.039 in.)
interlocking mechanism by shear was also observed. The
bottom section of the specimens experienced less deflection
likely due to the self-weight of the specimen.
Load (Bending) Perpendicular to the Bed Joints
Three specimens were tested under four point line of
load perpendicular to the bed joints. The dimensions of
the wallettes were 920 mm (36.22 in.) high (8 courses)
and 1,680 mm (66.14 in.) long (7 blocks). The positions
TMS Journal September 2007
of the outer bearings were 100 mm (3.94 in.) from the
edges the wallettes. The spacing of the inner bearing was
660 mm (25.98 in.). Rubber pad (packing) was introduced
between the bearing and the specimen. The specimens
were loaded under four point loading and lateral pressure
applied by means of a hydraulic jack supported by a rig
frame. Figure 15 shows the test specimen and experimental set up for line of load perpendicular to the bed joints.
Load cell and dial gauge were used to measure the load
and deflection respectively.
49
a) Elevation (AfterLoad Removal)
b) Plan (After Load Removal)
Figure 16—Failure Mode (Perpendicular to the bed joints)
Figure 17—Normalized load deflection curse (Flexure test)
The typical mode of failure of the wallettes tested
under load perpendicual to the bed joints was characterised
by the gradual rotation of the units about the horizontal
axis with load increase, leading into an opening/ splitting
of the head joints nearest to and between the loading points
(Figure 16a). This was accompanied by the formation of a
bending curve along the entire section (Figure 16b). Failure
of the interlocking mechanism by shear was also observed.
The bed joints practically remained very tightly closed. At
the point of failure the ultimate average lateral load was
about 12.4 kN (1,264.45 kp).
Analytical Model
Figure 17 shows the normalized load-deflection curves
of the wallets loaded parallel and perpendicular to the bed
joints. The loads were normalized with respect to the failure
load of the horizontal bending test (line of load perpendicular to the bed joints) and the deflections were normalized
with respect to the horizontal bending deflection at failure
50
load. The average failure load obtained from the flexural
tests was used to calculate the bending moment capacity of
the wallettes. Gross elastic section modulus was also calculated (ignoring the 4 mm (0.16 in.) tolerance between the
interlocking mechanisms). The ultimate transverse stress
was then calculated based on the relationship between the
bending moment and section modulus of the specimen. The
results are presented in Table 2.
Conventional masonry is not isotropic and therefore
does not provide the same resistance to bending in both
directions. The difference in resistance to bending when
spanning vertically (load parallel to bed joints) and horizontally (line of load perpendicular to bed joints) is defined
as the orthogonal ratio which is used primarily for the
calculation of bending moments in wall design. Similarly,
the test results suggest that dry-stack masonry does not provide the same resistance to bending in both directions. The
orthogonal ratio was calculated based on the ratio between
the ultimate flexural strength of dry stack masonry parallel
TMS Journal September 2007
Table 2. Result of Flexural Strength Calculations (1 kN = 101.97 kp and 1 N/mm2 = 145.037 psi)
Vertical Bending
Horizontal Bending
Mean failure load
Flexural strength
Mean failure load
Flexural strength
fkxparal-dry (N/mm²)
fkxperp-dry (kN)
(kN)
f kxperpend −dry (N/mm²)
1.5
0.03
12.4
and perpendicular to bed joints, as follows:
Orthogonal ratio of dry stack masonry =
µ dry =
f kxparal −dry
f kxperp−dry
=
0.03
= 0.2
0.15
(2)
Specimens tested under horizontal bending (line of
load perpendicular to the bed joint) performed better than
specimens tested under vertical bending (load parallel to
the bed joint). The flexural strength perpendicular to bed
joints was found to be about 0.15 N/mm² (21.76 psi), and
that parallel to bed joints was 0.03 N/mm² (4.35 psi) which
result in an orthogonal ratio of 0.2. In conventional masonry
this ratio is 1.2 to 3 [Curtin, Back, and Bray (1995)]. A
analytical relation for calculating the lateral load capacity
of the dry-stack system based on the flexural strength in
the stronger direction and the size of the panel is proposed,
as follows;


 h 2
h
P = f xdry ×10−3  48.38 − 6.32   −10.95
(3)
 L 
L


Where
P
= lateral load (peak), kN/m² (psi)
fxdry = flexural strength of the dry-stack masonry in
stronger direction, N/mm² (psi)
h, L = height and length of the panel respectively,
m (in.)
The result of analytical model plotted together with the
experimental results is shown in Figure 18. Comparison
between the experimental and computed results shows good
agreement. It should be noted that the above expression
may not be applicable for the blocks strength other than
9 MPa (1,305.34 psi).
DISCUSSIONS AND CONCLUSIONS
For conventional masonry, the ratio of panel strength
to the masonry unit strength is between 0.3 and 0.4 for inplane loading [Hendry (1991)]. This compares well with the
dry-stack system where the panel strength to unit strength
ratio is 0.35. The tests indicate a 65% increase in axial
load capacity when the blocks are bonded with mortar. The
reason for the lower compressive strength of the dry-stack
masonry is due to the mode of failure, which is predominantly in shear and splitting in the head and bed joints. The
contact area for transmitting load between units was approximately half of the total cross section area. The contact
area between the masonry units was 51% of the gross area
and this net area was used in stress calculation.
TMS Journal September 2007
0.15
Dry-stack wall panels under out of plane lateral load
showed significant similarities to conventional masonry such
as lifting of the starter course before failure, rotation of the
wall at the middle section, maximum deformation at the upper
half of the wall and the crack pattern. The crack patterns of the
dry-stack full scale walls under lateral load give “yield lines”
similar to that of laterally loaded concrete slabs. However,
due to the brittle nature of the masonry and dry-stacking of
the units, the similarity is geometrical only. The failure of the
panels was mainly due to the shear of the interlock during the
rotation of the units allowing the opening/splitting of both
the bed and head interlocking joints resulting in excessive
deflection of the wall. The dry-stack panels were found to be
stronger under horizontal bending likely due to the geometry
of the interlocking mechanism in the head joints.
Dry-stack masonry system has no standard code of
practice. Equation 3 could therefore be used as an initial
estimate of the lateral load capacity for dry-stack wall
simply supported on three edges, having h/L less than
2.0, provided the flexural strength of the masonry is first
established. The walls tested in this investigation are the
typical type of walls found in low-rise buildings, which is
the main focus of this study.
It is therefore the opinion of the author that this work
will be considered as preliminary approach towards development of standard procedures for studying the structural
behaviour of the dry-stack system. It is important also to
take into account the limited number of the specimens
tested in this investigation.
The salient conclusions of the study may be summarized as follows:
1. As in conventional masonry, test results show that
the wall panel strength of dry-stack wall under
axial compression is directly proportional to the
strength of the masonry units. The compressive
strength of the dry-stack wall was about 0.35 times
the compressive strength of the units.
2. Interlocking mechanism in the dry-stack units
assists alignment and stability of the wall under
axial loading.
3. The strength of dry-stack units does not make
significant difference in resistance to lateral load
as the interlocking keys and friction between units
govern the lateral load carrying capacity of the
dry-stack wall.
4. The crack pattern of dry-stack wall panel under
lateral load is very similar to the “yield line
51
pattern” in a typical laterally loaded reinforced
concrete slab.
5. Dry-stack wall panel under lateral load tends to lift
at the base before failure of the wall and rotates
about the middle section of the wall. Maximum
deformation occurs at the upper half of the wall
under lateral loading.
6. Vertical bending (load parallel to the bed joints)
failure occurs as “hinge-like rotation” along a bed
joint near the mid height of the wall.
7. Under horizontal bending (line of load perpendicular to the bed joints), the failure line occurs
along the dry-stack perpendicular joints between
the loading points.
8. The flexural strength perpendicular to bed joints (line
of load perpendicular to the bed joints) is about five
times more than parallel to the bed joints.
9. Like conventional masonry, dry-stack tested masonry is anisotropic.
10.Interlocking mechanism influence the flexural
strength of the masonry.
ACKNOWLEGMENT
This work was sponsored by the Hydraform-Africa
(Pty) Ltd. and the National Research Foundation. Their
support is gratefully acknowledged.
REFERENCE
Anand, K.B. and Ramamurthy, K., “Development of
Performance Evaluation of Interlocking-Block Masonry,”
Journal of Architectural Engineering, ASCE, Vol. 6, No.
2, pp. 45-51, 2000.
Agre´ment South Africa, Agreement certificate 69/237,
Pretoria, South Africa, 1996.
Curtin, W., Shaw, G., Back, J. and Bray, W., “Structural
Masonry Designers Manual,” 2nd edition, Blackwell Science Ltd, London, 1995
Drysdale, R.G. and Gazzola, E.A., “Strength and Deformation Properties of a Grounted, Drystacked, Interlocking,
Concrete Block System, Proc., 9th Int. Brick/ Block Masonry Conference, Deutsche Gesellschfur Mauerweksbau
e.V., Berlin, pp. 164-171, 1991.
Gallegos, H., “Mortarless Masonry: The Mecano System,
Int. J. Housing Sci. and its Applications, Vol. 12, No. 2,
pp. 145-157, 1988.
Gazzola, E.A. and Drysdale, R.G., Strength and Deformation Properties of Dry-stacked Surface Bonded Low-density
Block Masonry, Proc.5th Canadian Masonry Symposium,
Vancouver, Canada, pp. 609-618, 1989.
52
Hansen, K.F., “Strength and Deformation Capacity of Laterally Loaded Masonry”, Prco., 5th International Masonry
Conference, London, 1998.
Harris, H. G., Oh, K. and Hamid, A. A., “Development of
New Interlocking and Mortarless Block Masonry Units for
Efficient Building Systems”, Proceedings of The Sixth Canadian Masonry Symposium, Saskatoon, Canada, 1992.
Harris, H. G., Oh, K. and Hamid, A. A., “Development of
New Interlocking and Mortarless Block Masonry Units to
Improve the Earthquake Resistance of Masonry Construction,” Final Report to the National Science Foundation
Under Grant No. MSM-9102769, Department of Civil and
Architectural Engineering, Drexel University, Philadelphia,
USA, 1993.
Hatzinikolas, M., Elwi, A.E. and Lee, R., “Structural
Behaviour of Interlocking Masonry Block,” Proc., 4th
Canadian Masonry Symposium, Fredericton, Canada, pp.
225-239, 1986.
Hendry, A.W., “Structural Brickwork,” The Macmillan
Press Ltd, London, 1981.
Monk, C., “A Historical Survey and Analysis of the Compressive Strength of Brick Masonry,” Research Report
No. 12, Structural Clay Products Research Foundation,
Geneva, 1967.
Morsy, E., “An Investigation of Mortar Properties Influencing Brickwork Strength,” Ph.D. thesis, University of
Edinburgh, 1968.
NOTATIONS
fcu
= Unit strength of masonry.
fkxparal-dry = Ultimate flexural strength (load parallel to the
bed joints (Vertical bending).
fkxperp-dry = Ultimate flexural strength (line of load perpendicular to to the bed joints (horizontal bending)).
fpanel = Axial strength of wall panel.
fxdry
= Flexural strength in stronger direction.
h
= Height of masonry pane.
L
= Length of masonry wall panel.
L s
= Long span masonry pan.
Ms = Medium span masonry panel.
P
= Failure (peak) lateral load.
Ss
= Short span masonry panel.
t
= Thickness of masonry panel.
µ dry = Orthogonal ratio.
φ
= Material safety factor.
TMS Journal September 2007