MASONRY RESEARCH FOR EUROCODES
Miha Tomaževič, Vlatko Bosiljkov, Marjana Lutman
Slovenian National Building and Civil Engineering Institute,
Dimičeva 12, 1000 Ljubljana, SLOVENIA
SUMMARY
Some recent results of experimental research to support the requirements of Eurocodes for the
design and construction of innovative masonry structures in seismic zones are discussed.
Robustness of masonry units and type of masonry bond are important parameters which
define the behaviour of masonry walls when subjected to seismic loads. If the units are brittle
and the bond between units is not adequate, the known relationships between the strength and
ductility properties of masonry walls change significantly. In the design of masonry
structures, energy dissipation capacity is taken into account by force reduction, behaviour
factor q. The shaking table tests have shown that the behaviour factor does not depend only
on the masonry system, but also on the quality of materials and structural configuration.
Therefore, the values of behaviour factor q cannot be simply determined from ductility tests
of masonry walls.
1. INTRODUCTION AND SCOPE OF RECENT RESEARCH
Eurocode 6: Design of masonry structures - Part 1-1: Common rules for reinforced and
unreinforced masonry structures (EC 6) specifies the general design rules for masonry
structures [1]. In seismic zones additional rules, specified in Eurocode 8: Design of structures
for earthquake resistance - Part 1: General rules, seismic actions and rules for buildings (EC
8), apply, which in most cases overrule the requirements for the design of masonry structures
in non-seismic areas [2]. Most of these requirements are given in a general form, leaving the
National Annexes the task to provide specific values of parameters, depending on the
country’s specific conditions. It is not the aim of this contribution to discuss the contents of
Eurocodes. However, in order to better understand the recent need for research in masonry,
some of the most important parameters, which are expected to be defined in National
Annexes to Eurocode 8, are listed as follows:
−
−
−
−
type of masonry units with sufficient robustness
alternative classes of perpend (head) joints in masonry,
maximum value of ground acceleration for the use of unreinforced masonry,
the values of structural behaviour (force reduction) factor q for different masonry
structural systems.
These parameters cannot be determined unless systematic experimental research and
parametric studies are carried out. Whereas not so long ago there have been no problems with
robustness (or brittleness) of masonry units, and only fully filled head joints have been
allowed for the construction of masonry structures in seismic zones, innovation and
optimisation of technology of masonry construction introduced new types of masonry units
and construction technologies, which have so far not been regulated by the codes. Namely,
being pushed by competition on construction market, masonry industry improved the thermal
properties of masonry units and developed new, faster and cheaper technologies of
construction. As a result of such development, hollow masonry units with very thin shells and
webs are produced, and construction technologies are introduced where traditional head
joints, fully filled with mortar, are replaced by either ungrouted or partly grouted head joints
or mechanical interlocking between masonry units. The innovative proposals have been
developed in the countries not prone to seismic hazard. However, though not significantly
influencing the collapse mechanisms when subjected to gravity loads, these innovations
significantly influence the behaviour of masonry structures of all systems in seismic
conditions. It has been shown that they reduce the robustness of masonry units (due to thin
shells and webs) and homogeneity of masonry walls (due to masonry bond) as structural
elements.
New methods have been developed also for the verification of seismic resistance of masonry
structures. In order to take the advantage of these methods and optimise the resistance of
masonry buildings without risking the reduction of structural safety against earthquake loads,
experimental research is needed also to determine the values of structural behaviour factor q.
It is therefore of relevant importance to provide evidence that newly developed masonry units
and construction technologies are suitable for the use in earthquake prone regions. In fact,
several experimental research projects have been recently financed by masonry industry to
study the real situation, possibly relax the limitations given in the codes or propose
improvements in construction technology. The results of recent research, carried out at
Slovenian National Building and Civil Engineering Institute, and possible conclusions will be
summarized in this contribution.
2. ROBUSTNESS OF UNITS
According to EC 8, masonry units should have sufficient robustness in order to avoid local
brittle failure when subjected to seismic loads. The requirement is obvious, however, the
definition and criteria for robustness are lacking. In the previous version of EC 8, the
percentage of holes and the minimum thickness of shells in the case of hollow clay units have
been limited to 50 % and 15 mm, respectively. As past research and experiences indicate, this
requirement would guarantee solid behaviour of units and prevent unexpected behaviour of
structural walls due to local brittle failure of units when subjected to lateral seismic loads. In
the present version, however, the selection of suitable units according to grouping specified in
EC 6 is up to the National Annexes. Group 1 units are represented by solid or almost solid
units, whereas in the case of clay units belonging to Group 2, the amount of vertical holes is
limited to 55 % of the volume and the minimum thickness of shells and webs to 8 mm and 5
mm, respectively. This means that the shape of Group 2 clay units varies in a wide range from
almost solid to highly perforated hollow units. Without adequate quantification and testing it
is impossible to propose a selection.
Although the robustness of units has not yet been studied systematically, it recently proved to
be an extremely important parameter which governs the seismic behaviour of all systems of
masonry construction where hollow units are used. This has been the main conclusion made
on the basis of the analysis of test results obtained within the framework of two research
projects, recently carried out. The aim of the first project was to study the influence of the
amount of reinforcement on the lateral resistance and ductility of reinforced masonry walls
[3]. In the second case, however, the effect of different vertical bond types on lateral
resistance and ductility of plain masonry walls has been investigated. Recent research results
have clearly indicated that design recommendations, developed for the case of masonry walls
where no brittle behaviour of units is expected, have no validity in the case where the walls
fail in a non-ductile mode because of local brittle failure of units [4].
2.1 Robustness of units and effect of reinforcement
In the case of reinforced masonry sufficient anchorage and bond, as well as strength of the
units should be provided to utilize the tension capacity of reinforcing steel. To underline the
importance of robustness of masonry units in the case of reinforced masonry, the results of a
research project, where the suitability of a specially designed clay hollow masonry unit with
holes to accommodate vertical reinforcement has been investigated, will be summarized [3].
A series of tests of reinforced masonry walls, designed to fail in bending, has been carried
out. The shape and dimensions (175x290x190 mm - length x width x height) of the units used
for the construction of specimens were in conformity with the requirements of EC 6 for
Group 2 clay masonry units. The volume of the holes was 44 % of the gross volume of the
unit, whereas the thickness of shells and webs was 12 mm and 8 mm, respectively. The ratio
between the total thickness of webs in a section and the dimension of the unit across the same
section was 28 % which is more than required by EC 6 (16 %).
Two groups of four walls with different unit strength have been tested. The amount of vertical
reinforcement, placed in the holes at the ends of the walls and grouted with concrete, varied
as indicated in Table 1. Horizontal bed joint reinforcement was the same in all cases. In
addition to reinforced walls, two unreinforced walls of each unit strength class have been
made and tested as referential specimens.
Table 1. Designation and reinforcement of tested walls
Wall
Strength of
designation masonry
units
H1-0
6,2 MPa
H1-14
6,2 MPa
H1-20
6,2 MPa
H2-0
9 MPa
H2-20
9 MPa
H2-28
9 MPa
Vertical
steel
2 φ 14
2 φ 20
2 φ 20
2 φ 28
Vertical
reinforc.
ratio
2x0,056 %
2x0,115%
2x0,115%
2x0,225%
Horizontal Horizontal
steel
reinforc.
ratio
2 φ6/20 cm 0,099 %
2 φ6/20 cm 0,099 %
2 φ6/20 cm 0,099 %
2 φ6/20 cm 0,099 %
The walls have been 96 cm long, 140 cm high and 29 cm thick. General purpose mortar of the
strength class M2.5 has been used for the construction. The specimens have been built on r.c.
foundation blocks, fixed to the strong floor during the testing. In the case of reinforced walls,
deformed steel reinforcing bars, 14, 20, or 28 mm in diameter of the strength class M400,
have been placed vertically at the ends of the walls. The reinforcing bars have been anchored
into the foundation block at the bottom and into the bond-beam at the top of the wall. Stirrups,
placed in each horizontal mortar bed joint, have been bent around the vertical reinforcement
at the ends to improve the anchoring. Stirrups have been made of smooth reinforcing steel, 6
mm in diameter, of the strength class M250. Two specimens of each type have been tested.
The walls have been tested as vertical cantilevers, with foundation block fixed onto the strong
floor by means of steel bolts (Fig. 1). They have been subjected to constant vertical load and
cyclic lateral displacements, repeated three times at each displacement amplitude. Constant
compressive stress of 20 % of the compressive strength of masonry has been induced in the
walls’ horizontal section during the lateral resistance tests.
Figure 1. Typical testing arrangement for lateral resistance test of masonry walls
Figure 2. Lateral load - displacement
envelopes for walls H1:
0 - referential unreinforced wall
Figure 3. Lateral load - displacement
envelopes for walls H2:
0 - referential unreinforced wall
Experimentally obtained lateral resistance - displacement hysteresis envelopes are presented
in Figs. 2 and 3 for the walls type H1 and H2, respectively. In the case of reinforced
specimens, no difference in the resistance which can be attributed to different amounts of
vertical reinforcement, can be seen. There is, however, an improvement in lateral resistance
and ductility with regard to referential unreinforced walls of the same quality type, but the
improvement is attributed to the effect of the horizontal bed joint, and not vertical
reinforcement placed in the holes at the vertical edges of the walls. This can be concluded on
the basis of the measured strain in horizontal and vertical reinforcement.
Although the specimens have been designed to fail in bending, brittle local failure of units,
i.e. buckling and crushing of thin shells and webs, resulted into predominant ultimate shear
behaviour and collapse of the walls. Consequently, only a small part of the available tension
capacity of vertical reinforcing bars has been utilized. As can be seen in Table 2, where the
experimental values of lateral resistance Hexp are correlated with predicted values of flexural
capacity Hu,cal, the predicted values by far overestimated the actual resistance of the walls.
Typical crack patterns and details of damage at ultimate state are shown in Figs. 4 and 5. As
can be seen, local buckling and crushing of the shells of masonry units prevailed. On the basis
of the analysis of test results the conclusion can be made that in the case where local brittle
failure of hollow masonry units takes place, code specifications regarding the amount of steel
and calculation of section's capacity do not provide reliable results. In the best case, the
ductility and shear resistance capacity of such walls will be slightly improved due to
horizontal steel placed in the bed joints. However, vertical steel placed in the holes at the
edges of the walls will not improve the behaviour. Although the reinforced walls have been
designed “properly” by following the requirements and recommendations of the code, the
actual resistance will be only slightly greater than in the case where the walls were not
reinforced at all.
Figure 4. Brittle shear failure of wall
due to local brittle failure of units
Figure 5. Local buckling and crushing
of shells and webs of masonry units
In other words, the mathematical models used for the calculation of lateral in-plane resistance
of reinforced masonry walls' sections are reliable under the condition that sufficient bond
between mortar and steel as well as mortar and blocks is provided and that masonry units are
able to carry additional compression and shear which is transferred from the steel into the
units as a result of seismic loads. Yielding of steel at ultimate state is assumed in the
calculations of sectional capacity. However, if local brittle failure of masonry units takes
place and bond is degraded under cyclic loading before the tension capacity of reinforcing
steel is attained, the design value of resistance capacity of the reinforced masonry wall's
section will overestimate the actual resistance capacity. Although designed for earthquake
loads by taking into account code specifications, the actual seismic resistance of a masonry
structure under consideration could be lower than required and/or considered in seismic
resistance verification.
Table 2. Comparison between experimentally obtained Hexp and calculated
values of flexural capacity Hu,cal of tested walls
Wall
designation
H1-14
H1-20
H2-20
H2-28
h
(m)
1,60
1,60
1,58
1,58
Experimental
Hexp
Mexp
(kN)
(kNm)
80,9
129,5
78,2
125,2
109,1
172,4
103,2
163,0
Calculated
exp/cal
Hu,cal
Mu,cal
(kN)
(kNm)
87,6
140,2
0,92
118,3
189,3
0,66
136,6
215,9
0,80
173,3
273,8
0,60
2.2. Robustness of units and bonding patterns
Traditionally, only the construction of masonry walls with fully grouted (filled) head joints
has been allowed in seismic zones. By definition given in EC 6, perpend joints can be
considered to be filled if mortar is provided to the full height of the joint over a minimum of
40 % of the width of the unit. In the new draft of EC 8, however, new techniques of bonding
the masonry units vertically are also allowed to be used in the construction of buildings in
seismic zones, such as fully grouted, ungrouted, and ungrouted joints with mechanical
interlocking between the units. The decision, which of the three classes of perpend joints will
be allowed to be used in a country, is left to the National Annex of each individual country.
In order to make possible such decision and provide the necessary basic experimentla data, a
research project has been proposed to investigate the influence of different systems of filling
the perpend joints on the seismic behaviour of masonry walls. Four different types of filling
the perpend joints have been studied (Fig. 6):
•
•
•
•
Fully filled perpend joint (walls series BN - referential)
Dry, unfilled perpend joint (walls series BG)
Partly filled perpend joints with mortar in the pockets (walls series BP)
Dry, grove and tongue perpend joint (walls series BZ)
The walls, designed to fail in shear, have been 125 cm long, 150 cm high and 30 cm thick.
Grade M10 clay hollow units with dimensions 245x300x240 mm (length x width x height),
with 12 mm thick shells, 8 mm thick webs and 50 % of holes per volume of the unit (Group 2
units according to EC 6) have been used, same in all cases except for the head joint face (Fig.
7). Grade M5 general purpose mortar has been used for the construction of specimens, built
on r.c. foundation blocks and tested in the same way as described in 2.1 and shown in Fig. 1.
Three specimens of each type have been tested under constant vertical load, which induced
vertical stresses in the horizontal section amounting to 1/3 of the compressive strength of the
masonry [4].
Figure 6. Schematic presentation of the investigated types of perpend joints
Figure 7. Masonry units for the construction of walls of series BN and BG
All walls failed in shear, as expected. However, brittle local failure of units, i.e. buckling and
crushing of thin shells and webs was the predominant phenomenon which determined the
failure mode. The walls failed in a non-ductile, brittle mode, soon after the maximum
resistance has been attained. The test results are summarized in Tables 3 and 4. In Table 3,
mechanical characteristics of the tested types of walls, such as compressive f and tensile
strength ft, as well as modulus of elasticity E and shear modulus G, are evaluated.
Table 3. Mechanical characteristics of the tested types of masonry walls
Type
f
(MPa)
ft
(MPa)
E
(MPa)
G
(MPa)
ft/f
G/E
BN
BG
BP
BZ
4,13
4,31
6,28
6,24
0,17
0,19
0,22
0,20
3088
3302
4815
5548
330
354
320
367
0,041
0,044
0,035
0,033
0,107
0,107
0,066
0,066
It can be noticed that the ratio between the tensile and compressive strength of masonry ft/f of
all tested types is low and did not exceed 4 % of the value of compressive strength (usually 6
-8 %). The ratio between the values of shear modulus G, evaluated on the basis of the
effective stiffness of the walls measured during lateral resistance tests, and modulus of
elasticity E, evaluated on the basis of compression tests, did not exceed 0.11. The ratio
decreased with the increased compressive strength of masonry.
In Table 4, the lateral resistance and displacement capacity indicators are given for the tested
types of wall, defined as the ratios between the lateral resistance and displacement values at
different limit states, such as crack (dcr, Hcr), maximum resistance (dHmax, Hmax), and ultimate
state limit (du, Hu).
Table 4. Lateral resistance and ductility capacity indicators of the tested types of walls
Series
Hcr/Hmax
Hdu/Hmax
dcr/dHmax
du/dHmax
du/dcr
BN
BG
BP
BZ
0,98
1,00
0,97
0,93
0,44
0,66
0,45
0,34
1,00
1,00
0,84
0,68
2,15
1,13
1,73
1,75
2,15
1,13
2,06
2,57
A conclusion can be made that no significant influence of different types of filling the
perpend joints can be observed on the lateral resistance and displacement capacity of the
tested walls. The ratio between the lateral load acting on the walls at the initiation of diagonal
cracking and maximum resistance Hcr/Hmax was close to 1.00. This means that the occurrence
of diagonal shear cracks in masonry characterizes the attainment of lateral resistance of the
walls. In the particular case studied, once the diagonal cracks occurred the resistance of the
walls at increased imposed displacements amplitudes started to degrade. Relatively large
resistance degradation has been observed in all cases, i.e. small values of Hdu/Hmax ratio have
been obtained despite the relatively small ultimate displacements, hence indicating the brittle
character of the behaviour of the tested walls at shear failure. The displacement capacity of
the tested walls in terms of displacement capacity indicators is small, below the expected
values for unreinforced masonry walls.
The actual values of mechanical properties of masonry, obtained by testing, are compared
with the values, calculated by considering EC 6 equations and procedures, in Table 5.
Table 5. Correlation between experimental and EC 6 predicted values of shear strength and
shear modulus of masonry
Series
fvexp
(MPa)
Gexp
(MPa)
fvkEC6
(MPa)
GEC6
(MPa)
fvexp/fvkEC6
Gexp/GEC6
BN
BG
BP
BZ
0,39
0,38
0,43
0,39
330
354
320
367
0,65
0,45
0,65
0,65
1464
1464
1464
1464
0,60
0,84
0,66
0,60
0,22
0,24
0,22
0,25
It can be seen that the code predicted values of mechanical properties which define the
seismic resistance of walls at shear failure (shear strength fvk and shear modulus G)
overestimate the actual, experimentally obtained values in almost all cases.
Taking into consideration the observed resistance and displacement capacity as well as
mechanical properties of the tested wall types, it can be concluded that in all cases the
behaviour of the specimens subjected to cyclic lateral loading has been governed by the
premature local brittle failure of units. Consequently, the failure of the walls subjected to
cyclic lateral loading occurred at a stage where the type of filling the vertical, head joints did
not yet influence the behaviour of the walls. The failure mechanisms in all cases depended on
the characteristics of masonry units. Against expectations, they did not depend on the way of
construction of the wall specimens, i.e. the type of masonry bond or the way of how the head
joints have been filled.
Therefore, no firm general conclusion can be made as regards the influence of types of
masonry bond on the seismic behaviour of masonry walls. Additional research is needed, and
masonry units with sufficient robustness to avoid local brittle failure should be used for the
construction of test specimens in order to ensure that the masonry bond and not the units
influence the behaviour of the walls when subjected to cyclic lateral load.
3. STRUCTURAL BEHAVIOUR FACTOR q
With exception of the so called simple buildings, i.e. buildings which fulfill severe limitations
regarding the height, structural configuration and quality of materials, the stability of masonry
structures for vertical and seismic loads should be verified by calculation. As is the case of
other types of structures, the seismic resistant design of masonry structures of all systems by
EC 8 is based on:
• no collapse requirement, and
• damage limitation requirement.
Taking into account the regularity of masonry buildings whose response is not significantly
affected by contribution from higher modes of vibration (e.g. [6] and [79]), lateral force
method of analysis will provide adequate results. Following this method, the seismic base
shear force Fb for each horizontal direction in which the building is analysed, is determined
by EC 8 as follows:
Fb = Sd(T1) m λ,
(1)
where:
Sd(T1) = the ordinate of the design spectrum at period T1,
T1 = fundamental period of vibration of the building for lateral motion in the direction
considered,
m = total mass of the building above the foundation or above the top of a rigid basement
λ = correction factor, accounting for the fact that in building with at least three storeys and
translational degrees of freedom in each horizontal direction, the effective modal mass of
the 1st mode is smaller than the total building mass.
Masonry buildings are rigid structures with natural periods of vibration ranging between
periods where the EC 8 response spectrum is flat. Therefore, the ordinate of the design
spectrum for masonry buildings can be determined by:
Sd (T) = a g S η
2,5
,
q
(2)
where:
ag = design ground acceleration on type A ground (rock or rock-like formation),
S = soil factor,
η = damping correction factor (η = 1 for 5% viscous damping),
q = structural behaviour factor.
As specified in EC 8, the capacity of structural system to resist seismic actions in the nonlinear range generally permits the design for forces smaller than those corresponding to a
linear elastic response. To avoid explicit inelastic structural analysis in design, the capacity of
the structure to dissipate energy through mainly ductile behaviour of its elements and other
mechanisms is taken into account by performing an elastic analysis based on a response
spectrum reduced with respect to the elastic one by introducing the behaviour factor q.
In a qualitative and simplified way, the definition of behaviour factor q is explained in Figure
8, where the seismic response curve of an actual structure, idealized as a linear elastic perfectly plastic envelope, is compared with the response of a perfectly elastic structure
having the same initial elastic stiffness characteristics.
Figure 8. Definition of structural behaviour factor q
As a result of the energy dissipation capacity of the actual structure, expressed by the global
ductility factor µu = du/de, there is no need for the structure to be designed for strength, i.e. for
the expected elastic load He. The structure is designed for the ultimate design load Hdu and the
ratio between the two is called the behaviour factor q:
q = He/Hdu.
(3a)
In other words, behaviour factor q is an approximation of the ratio between the seismic force
which the structure would experience if its response is completely elastic and minimum
seismic force which may be used in the design with a conventional elastic model, still
ensuring a satisfactory response of the structure. Following the definition in Fig. 1, structural
behaviour factor can be also expressed in terms of the global ductility factor µu = du/de as
follows:
q = (2 µu - 1)1/2.
(3b)
A range of values of q factor for different systems of masonry construction is proposed in EC
8:
•
•
•
for unreinforced masonry: q = 1.5 - 2.5
for confined masonry: q = 2.0 - 3.0
for reinforced masonry: q = 2.5 - 3.0
Since the proposed values have limited experimental background. However, as some
preliminary testing indicated that the proposed values are conservative [7], a research project
has been recently carried out to investigate the seismic behaviour and determine the range of
values of structural behaviour factor for selected types of unreinforced masonry buildings [8].
3.1. Experimental program and description of tests
Six models representing buildings with two different structural configurations and two
different types of masonry materials have been tested on a simple uni-directional seismic
simulator, a two-storey terraced house with main structural walls orthogonal to seismic
motion (models M1 - Fig. 9) and a three-storey apartment house with uniformly distributed
structural walls in both directions (models M2 - Fig. 10). Four models of the first and two
models of the second type have been tested. In the case of the terraced house, two models
have been built as either partly or completely confined masonry structures (Table 6).
Table 6. Shaking-table tests - description of tested models
Designation
Type
Material
Remark
M1-1
M1-2
M1-1c
M1-1d
M2-1
M2-2
Terraced house
Terraced house
Terraced house
Terraced house
Apartment house
Apartment house
Calcium silicate
Hollow clay unit
Calcium silicate
Calcium silicate
Calcium silicate
Hollow clay unit
no confinement
no confinement
confined staircase walls
fully confined walls
no confinement
no confinement
Because of the limited capacity of earthquake simulator installed at ZAG, models have been
built in a 1:5 scale from special model materials designed to fulfill the requirements of
complete model similitude. The correlation between the model and prototype characteristics
of masonry has been verified by testing a series of model and prototype-size walls under
compression as well as under combination of constant compression and cyclic lateral loading.
Nevertheless, it should be borne in mind, that by testing small scale masonry models, only the
global behaviour and mechanism of the buildings’ behaviour can be adequately simulated,
and not the behaviour of structural details.
The models have been built on foundation slabs which have been fixed before the tests to the
steel platform of the shaking table by means of bolts. In order to fulfill the requirements for
similitude of dynamic behaviour, additional masses (steel bricks) have been fixed to each
floor slab to simulate the effect of the live load. The models have been instrumented with
displacement meters (LVDTs) and accelerometers fixed on the model at both edges and
centre of each floor slab.
The north-south component of earthquake acceleration record obtained at Petrovac during the
April 15, 1979, earthquake in Montenegro, with peak ground acceleration of 0,43 g has been
used to drive the shaking-table. The intensity of shaking was controlled by adjusting the
maximum amplitude of the shaking-table displacement, obtained by numerical integration of
the accelerogram used as the input in each successive test run, scaled according to the laws of
model similitude. The analysis of shaking-table motion, carried out during one of our
previous studies [9] has shown that the absolute acceleration spectra of the shaking-table
motion, normalized with regard to maximum acceleration, are in good agreement with one of
the previous versions of the EC 8 response spectrum.
All models have been tested by subjecting them to step-wise increasing intensity of the
shaking-table motion in each subsequent test run until models’ final collapse. During the
tests, displacement and acceleration responses of the models at each floor have been
measured. The behaviour of the models during testing has bee video-taped for further analysis
of damage propagation. After each test run, the models have been inspected for damage, and
the cracks have been marked and photographed. Also, the changes in dynamic properties of
the models have been determined by analysing the records of free vibrations obtained by
hitting the top slab of the model with hammer.
3.2. Test results
All models failed in shear, as expected. Regardless to the structural type and configuration,
shear cracks developed in structural walls in the direction of seismic motion, subsequently
leading to stiffness and strength degradation and final collapse of the models. The
unreinforced terraced house model M1-1 and apartment house model M2-1, made of
materials simulating calcium silicate masonry units collapsed immediately after the first
damage occurred, whereas respective models M1-2 and M2-2, made of materials simulating
hollow clay units, though damaged, withstood additional shaking before collapse. The
behaviour of partly and fully confined terraced house models M1-1c and M1-1d was
significantly improved. Typical damage to the models just before collapse is shown in Figs. 9
and 10 for a terraced and apartment house model, respectively.
Figure 9. Confined terraced house
model M1-1c just before collapse
Figure 10. Apartment house model
M2-2 just before collapse
As can be seen, the damage to structural walls was concentrated in the first storey, so that
typical shear type mechanism of seismic behaviour prevailed. Very little damage to the walls
in the upper storeys has been observed at the moment of collapse in all cases, including
confined terraced house models M1-1c and M1-1d. As a result of this mechanism, relative
displacements of the upper storeys in the non-linear range of vibration were very small
compared to the first storey drift. As the deformations and the amount of damage were small,
the amount of dissipated, hysteretic energy in the upper storeys did not exceed 5 % of the
energy dissipated in the first storey. Taking this into consideration, the conclusion can be
made that the resistance envelope of the first storey determines the seismic behaviour of the
tested structures.
On the basis of the recorded displacement and acceleration response time histories and taking
into account the masses of the models, concentrated at each floor level, the maximum values
of the base shear developed in the models during the individual phases of testing, have been
calculated. The values have been expressed in a non-dimensional form in terms of the base
shear coefficient (BSC), which is the ratio between the base shear resisted and the weight of
the model. The values are plotted against the first storey rotation angle (the ratio between the
relative storey displacement and storey height), hence obtaining the lateral resistance displacement envelopes of a critical storey in a non-dimensional form (Fig. 11).
2
0,8
Models M1
1,6
Models M2
0,6
Model M1-2
1,2
BSC
Model M1-1c
BSC
Model M1-1d
Model M1-1
0,8
0,4
Model M2-1
0,2
0,4
Model M2-2
0
0
0,01
0,02
0,03
0,04
Rotation angle
0,05
0,0
0
0,005
0,01
0,015
0,02
Rotation angle
Figure 11. Experimentally obtained base shear coefficient - storey rotation angle relationships
As can be seen, the values of storey rotation angle where the stiffness of the models has
significantly changed as a result of the damage occurring to structural walls (damage limit),
are very close in all cases. The values of 0,25 % have been measured in the case of the
terraced house models M1 and the values of 0,3% in the case of the apartment house models
M2. It can be also seen that the damage limit values of storey rotation angle coincide or are
very close to the values where the maximum resistance has been attained.
Regarding the influence of masonry materials on the seismic behaviour of the tested
buildings, it can be seen that the models of both, terraced house and apartment house
structural type, made of model materials simulating calcium silicate masonry units (models
M1-1 and M2-1) exhibited substantially more brittle behaviour than the models of the same
type, but made of model materials simulating hollow clay units. However, there has been not
much difference observed as regards the resistance. The confinement of structural walls with
vertical r.c. confining elements in the case of the terraced house models M1-1c and M1-1d
proved to be a successful measure to improve the seismic behaviour of the terraced house as
regards both lateral resistance and displacement capacity.
3.3 Structural behaviour factor q
In order to evaluate the values of behaviour factor q, the basic definition given in EC 8 and
the previously explained simple philosophy have been followed. Typical examples are shown
in Figs. 12 and 13 for the terraced and apartment house models, respectively.
The values of behaviour factors q, resulting from experimental envelopes, are presented in
Table 7. In the calculations of the elastic response of the models, effective stiffness of the
model structure, i.e. the measured stiffness at the occurrence of the first significant damage to
structural walls (damage limit), has been taken into account, and not the initial stiffness of the
model, measured before the shaking-table tests. EAVEK, a commercial computer program for
seismic analysis of multi-storey buildings, has been used. Following the definition of
behaviour factor q, the response of the elastic structure subjected to shaking-table motion
during the testing phase in which the maximum resistance of the model has been attained, has
been calculated.
1,0
3
M1-1
0,8
M1-1d
2,5
2
BSC
BSC
0,6
R010
0,4
R050
R005
Experimental
Elastic response
Idealized
R075
1
R050
R025
R150A
0,5
0,2
R025
R010
R005
0
0,0
0
0
0,01
0,02
0,03
Experimental
Elastic response
Idealized
R100
1,5
0,01
0,04
0,02
0,03
0,04
Rotation angle
Rotation angle
Figure 12. Base shear coefficient - storey rotation angle relationships obtained for
plain and confined terraced house models
1,2
1,2
M2-1
1
1
0,8
R125
0,6
BSC
BSC
0,8
Experimental
Elastic response
Idealized
R100
R125
R075
0,4
0
0
R100
R075
R050
0,4
Experimental
Elastic response
Idealized
R050
R025
R010
R005
0,2
0,6
0,01
0,02
Rotation angle
0,03
R100
R100
R025
0,2
R010
R005
0
0,04
M2-2
0
0,01
0,02
0,03
0,04
Rotation angle
Figure 13. Base shear coefficient - storey rotation angle relationships
obtained for apartment house models
The elastic response in terms of maximum elastic base shear has been compared with the
maximum experimental base shear value BSCmax as well as with the value BSCu obtained by
the idealization of experimental resistance envelope, following the definition given in Eq. 3a.
As can be seen in the diagrams in Figs. 12 and 13, the experimentally obtained envelopes
have been idealized as bilinear elastic-plastic relationships. In the cases where no sudden
collapse has occurred, 20 % of strength degradation has been allowed as a measure to
evaluate the idealized global ductility µu of the structure without risking collapse. However, it
has to be noted that substantial damage to structural walls of the models has occurred at that
stage. In order to fulfill also the “damage limitation” requirement, only part of the available
displacement capacity has been taken into account for the evaluation of behaviour factor q on
the basis of global ductility of the structure (Eq. 3b), limited by the displacement value where
severe damage to structural walls occurs. This value has been arbitrarily chosen to be 3-times
the value of storey rotation at the damage limit Φu = 3 Φdam. Typical damage to structural
walls at the maximum permissible storey rotation angle (ultimate damage limit) is shown in
Fig. 14.
Table 7. Values of structural behaviour factor q evaluated from experiments
Model
M1-1
M1-2
M1-1c
M1-1d
M2-1
M2-2
q = BSCe/BSCmax
1,34
1,84
2,44
1,63
1,55
1,74
q = BSCe/BSCu
1,53
2,06
2,56
1,91
1,91
1,85
q = (2 µu - 1)1/2
2,09
2,20
2,99
2,88
2,61
2,84
Fig. 14. Typical damage to structural walls at ultimate
damage limit (model M1-1c)
It can be seen that, generally speaking, the evaluation of behaviour factor q on the basis of the
observed ductility capacity of the models resulted into higher values than the simple
correlation between the values of the theoretical elastic and observed base shear responses. It
can be also seen that despite the differences observed in the behaviour of unreinforced
masonry models during shaking-table tests (brittle behaviour of models M1-1 and M2-1 made
of calcium silicate units against “ductile” behaviour of models M1-2 and M2-2 made of
hollow clay units - see Fig. 11), the values of behaviour factor q, evaluated on the basis of
simple definition, are of the same order of magnitude for the cases of both, terraced and
apartment house structural types. In the case of the terraced house models with confined
structural walls (models M1-1c and M1-1d), however, the observed improved behaviour
resulted also in the increased evaluated values of behaviour factor q.
Although the resistance and displacement capacity can be used as a measure of energy
dissipation capacity, taking advantage of the measured response data, energy dissipation
capacity of the tested models has been calculated on the basis of the measured storey shear storey drift (relative storey displacement) hysteresis loops. The results of calculations are
given in Table 8, where for each of the tested models the input energy, induced to the system
during shaking by hydraulic actuator and dissipated hysteretic energy, are presented.
Cumulative values of input and dissipated hysteretic energy from the beginning to the end of
the shaking-table tests, are given in the table, as well as the ratio between both..
Table 8. Relationships between the cumulative input and dissipated
hysteretic energy at the end of shaking-table tests
Model
M1-2
M1-1c
M1-1d
M2-1
M2-2
Input energy
Einp (Nm)
2710
1866
4813
565
2352
Dissipated hysteretic energy
Ehys (Nm)
1267
637
1778
94
750
Ehys/Einp
0,47
0,34
0,37
0,17
0,32
In order to make conclusions on the basis of data presented in Table 8, one has to take into
consideration that, although both parameters are a function of structural response, input
energy (energy demand) is a function of masses of the structure, ground (shaking-table)
acceleration time history and velocity response of the structure, whereas the amount of
dissipated hysteretic energy is determined mainly by damage propagation mechanism.
Therefore, only the comparison of data obtained for the models of the same structural
configuration is reasonable.
In this regard, the differences in the observed behaviour of apartment house models M2-1 and
M2-2, made of calcium silicate and hollow clay units, respectively, can be also explained by
the differences in energy dissipation capacity. Model M2-2 needed 4-times more input energy
Einp than model M2-1 to cause collapse. At the same time, energy dissipation capacity Ehys of
model M2-2 was almost 8-times greater than energy dissipation capacity of model M2-1. As a
result, the difference in Ehys/Einp ratio is also significant. It is assumed that similar observation
could have been made for the case of unreinforced terraced house models M1-1 and M1-2, if
the response records of model M1-1 were available in a form suitable for the analysis.
As the data for terraced house model M1-1 are missing, it is not possible to evaluate the effect
of partial and complete confinement of structural walls with regard to referential model in
terms of energy dissipation capacity. However, the difference between the partly and fully
confined models M1-1c and M1-1d can be clearly seen. 2.6-times more input energy has been
needed in the case of fully confined model M1-1d to cause collapse than in the case of partly
confined model M1-1c, and 2,8-times more energy has been dissipated. Obviously, this
resulted in almost the same Ehys/Einp ratios in both cases.
The decision about which values of behaviour factor q to recommend for the design of the
tested types of buildings is therefore not a simple one. Namely, following the simple
definition of behaviour factor according to EC 8, the differences in the assessed values for
unreinforced buildings of both structural configuration types are not significant (Table 7).
However, significant differences have been observed in the behaviour of the models of both
structural types, made of the calcium silicate units on the one hand and those made of hollow
clay units on the other. As can be seen by comparing the experimentally obtained resistance
envelopes, the behaviour of the models, made of hollow clay units was significantly more
ductile than the behaviour of the models made of calcium silicate units. The differences can
be also seen by comparing the calculated input and dissipated hysteretic energy balance
(Table 8). Obviously these observations need to be taken into consideration when making the
final proposal. However, it can be concluded that the obtained values are well within the
range of EC 8 proposed values for unreinforced and confined masonry structural systems.
It has to be underlined, that the actual resistance envelopes of the tested building models have
been considered in the evaluation of the values of structural behaviour factor. Since the basic
requirements for adequate resistance and damage limitation have been explicitly taken into
account, these values represent the minimum values of structural behaviour factor q which
may be considered in the case that accurate mechanism models and pushover analysis are
used for the verification of seismic resistance of the structures under consideration. However,
in the case of most practical design methods, the values of q factor could be further increased
as a results of the structural overstrength. Parametric studies are needed to estimate the
expected ranges of ovestrength resulting from the most widely used methods of calculation.
4. CONCLUSIONS AND PLANS FOR FUTURE RESEARCH
As the recent experimental investigations indicated, the development and innovation in
masonry brought to the front the robustness of masonry units as one of the basic parameters,
which determine the seismic behaviour of masonry walls and structures. In the case of the
local brittle failure of masonry units, the failure mechanisms of the walls subjected to seismic
loads change, so that, because of the changed basic relationships, the validity of most
practically used calculation procedures for seismic resistance verification of unreinforced and
reinforced masonry structures becomes questionnable.
This has been also recognized in the new draft of EC 8, where the requirement that masonry
units, used for the construction of masonry structures in seismic zones, should have sufficient
robustness to avoid local brittle failure, is given. However, no quantitative criteria to fulfill
such requirement are given in EC 8. Therefore, additional experimental research is needed to
provide the basis and criteria for the classification of hollow masonry units regarding the
robustness. It is planned that, as a result of a recently initiated research project, such criteria
will be determined and a testing method for simple evaluation of robustness will be
developed. The influence of mortar strength and the way of laying the units will be also
studied, along with the influence of different types (classes) of head joints (fully and partly
grouted, ungrouted, mechanical interlocking) on the homogeneity of structural walls, built
with masonry units which will comply with the criteria for sufficient robustness.
As regards the energy disipation capacity of masonry structures and EC 8 suggested values of
structural behaviour factors q, the experimental study indicated that the ranges of values of
factor q proposed in EC 8 for different masonry systems, are adequate. However, the study
also indicated that the values depend not only on the system of construction, but also on the
properties of masonry materials and structural configuration of the building under
consideration. Therefore, experimental research is needed for the assessment of a particular
value for a particular structural type within the recommended range of values. However,
although such tests are helpful, the values of behaviour factor q cannot be assessed by means
of only ductility tests of structural walls.
It has to be also underlined that the values recommended on the basis of the recent
experimental study (q = 1,5 for unreinforced houses of both tested types and q = 2,0 for
confined masonry houses) are valid for regular masonry structures, with uniformly distributed
walls in both orthogonal directions and along the height. In the particular case studied, the
structural configuration of the apartment house buildings fulfilled the requirements for
regularity, whereas the configuration of the terraced house building was on the edge of the
acceptable limit. Requirements should be provided for minimum amount and position of
structural walls for the case that the behaviour factor q = 1,5 is used in the design of terraced
house type of buildings.
5. ACKNOWLEDGEMENTS
This paper discusses the results of several research projects, financed by the Ministry of
Higher Education, Science and Technology of the Republic of Slovenia, Chamber of
Commerce and Industry of the Republic Slovenia, and associations of brick masonry
producers from Slovenia (Wienerbeger Opekarna Ormož and Goriške opekarne), Austria
(Verband Österreichischer Ziegelwerke), Germany (Deutsche Gesellschaft für
Mauerwerksbau), Italy (Associazione Nationale Degli Industriali dei Laterizi) and
Switzerland (Verband Schweizerische Ziegelindustrie). Their financial and professional
contribution to the outcome of the projects is gratefully acknowledged.
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