Engineering Structures 206 (2020) 110166
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Engineering Structures
journal homepage: www.elsevier.com/locate/engstruct
Large-scale testing of low damage rocking Cross Laminated Timber (CLT)
wall panels with friction dampers
T
Ashkan Hashemi , Pierre Quenneville
⁎
Department of Civil and Environmental Engineering, Faculty of Engineering, The University of Auckland, Private bag 92019, Auckland 1142, New Zealand
ARTICLE INFO
ABSTRACT
Keywords:
Rocking walls
Cross Laminated Timber
Friction
Self-centring
Low damage
Resilience
The use of controlled ductile rocking shear walls with low damage connections can be an efficient alternative to
traditional high damage design in order to mitigate the earthquake induced damage. Nowadays there is an
increased demand for seismic low damage systems that can offer a high resistance level against the severe
seismic events, allowing the buildings to be swiftly returned to service. This paper introduces a friction-based
energy dissipation hold-down connection for rocking Cross Laminated Timber (CLT) structures with potential for
re-centring behaviour. The results of component experimental tests on the device and results of large-scale tests
on a rocking CLT wall with the proposed hold-down connectors have shown that the proposed system has the
potential to be used in earthquake-resistant low- to mid-rise CLT structures. Furthermore, capacity prediction
equations are provided for this system that are verified by comparing the analytical data with the experimental
results.
1. Introduction
In recent years, Cross Laminated Timber (CLT) has been widely used
for different types of buildings such as offices, commercial buildings,
public buildings and multi-story residential complexes. These panels,
made of multiple layers of timber boards arranged in a crosswise configuration [1,2], are dimensionally stable and can offer mechanical
characteristics similar to those of pre-cast reinforced concrete panels.
For conventional CLT structures (panelised structures), traditional steel
connectors with dowel type fasteners such as nails and screws are extensively being used. Despite the good seismic performance of CLT
structures in in high seismically active regions, the experienced damage
due to inelastic deformation of the steel brackets is difficult and costly
to repair.
One of the most extensive experimental research about the seismic
performance of CLT structures to date has been conducted within the
SOFIE project [3]. That project included quasi-static tests on a singlestory CLT building with different layouts, shake table tests on a threestory CLT building and a series of full-scale shake table tests on a sevenstory CLT building. The results confirmed that the CLT structures with
traditional connections are relatively stiff and could survive the design
level events (although significant damage was observed in some connections). Nevertheless, some of the steel connections (such as nailed
hold-downs and nailed shear brackets) yielded in bending and some
⁎
withdrew from the timber elements. More importantly, high response
accelerations (mainly in the upper levels) with a maximum of 3.8 g
were recorded. Accelerations this high have the potential to seriously
harm the occupants, and thus as an important outcome of that research,
it was recommended to consider a solution to reduce them.
Popovski et al. investigated the seismic response of the CLT wall
panels with various arrangements and connection layouts for application in platform structures [4,5]. The data would seem to suggest that
these walls have satisfactory lateral resistance when nails or slender
screws are used together with steel brackets. Moreover, the use of holddowns with nails at the corners of the walls was proven to improve the
resistance to overturning from the lateral forces. That can be attributed
to the increased moment lever arm for the wall panels.
Gavric et al. experimentally investigated the cyclic behaviour of
single and coupled CLT walls with different connections [6]. The test
results suggested that the layout and design of the connections govern
the overall behaviour of the wall. While in-plane deformations of the
panels were almost negligible, inelastic deformations in the connection
parts lead to local failures in the system.
Popovski et al. conducted a series of full scale quasi-static tests on a
two-story platform house [7]. No global instability was observed not
even at the maximum design lateral force. Regardless of the rigid
connections between the floors and walls, rocking movement of the
wall panels was not totally restricted by the floor above.
Corresponding author.
E-mail addresses: ahas439@aucklanduni.ac.nz (A. Hashemi), p.quenneville@auckland.ac.nz (P. Quenneville).
https://doi.org/10.1016/j.engstruct.2020.110166
Received 15 August 2019; Received in revised form 16 December 2019; Accepted 1 January 2020
0141-0296/ © 2020 Elsevier Ltd. All rights reserved.
Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
Yasumura et al. studied the mechanical performance of low-rise CLT
platform building with large and small panels subjected to reversed
cyclic lateral loads [8]. It was concluded that in the buildings with
small panels, rotation of the panels was the major cause of the total
deformation of the building. They also proposed a numerical model to
predict the seismic behaviour of such structures based on the connectors in use. Blomgren et al. [9] performed a series of shake table
tests on a full-scale, two-storey mass timber building with rocking CLT
shear walls. The shear walls were equipped with sacrificial components
(fuses). The test results showed that the system could meet the design
performance indexes and the damage was limited to the designated
components that could quickly be replaced after the event.
Regardless of the satisfactory seismic resistance of the abovementioned CLT structures, in almost all cases, the connections suffered
from large plastic deformations at the end of the earthquake or the end
of the cyclic test [7]. This means that many of these connectors would
need to be repaired or replaced after a major seismic event which can
be considerably costly or even not practical.
The use of friction devices for mitigating the seismic damage dates
back to 1980s where Pall et al. [8] introduced them to be used in reinforced concrete panels and steel braced frames. Later on, Popov et al.
[10] and Clifton et al. [11] proposed the use of friction connections for
steel moment resisting frames. For timber structures, Filiatrault [12]
investigated the application of friction damping devices at the corners
of traditional sheathed timber shear walls that demonstrated a promising improvement in the hysteretic behaviour of the wall system.
Bora et al. introduced the use for friction-based energy dissipated holddowns for rocking concrete walls [13]. Loo et al. [14] investigated the
application of flat plates slip friction connections as a replacement to
traditional hold-downs for timber Laminated Veneer Lumber (LVL)
walls. Their experiments showed a significantly enhanced seismic performance compared to traditional systems in terms of stability and
hysteretic damping. Nonetheless, it was revealed that a self-centring
behaviour can only be achieved by applying additional gravity loads to
the rocking wall [15]. Hashemi et al. [16–18] expanded the concept of
slip friction connections to the CLT coupled walls and hybrid timbersteel core walls. It was shown that despite the reliable low damage
seismic performance of the proposed systems, an additional mechanism
and/or vertical gravity loads is required to provide a self-centring behaviour. Later on, Hashemi et al. developed seismic solutions for timber
structures using resilient self-centring connections with innovative
friction dampers [19–21].
This paper proposes a friction-based energy dissipation hold-down
for rocking CLT shear walls that can be tuned to significantly improve
the re-centring capacity of the wall system while absorbing the seismic
input energy. The joint component test results and large scale experimental results showed the efficiency of the system for the potential use
in CLT structures when a low damage solution is required.
the linear pre-stressed die springs can be in contact with it using stiffener plates. A rod is considered in the middle of each spring that is
anchored to the foundation.
At the start of loading, the applied load to the connection needs to
increase to the level that overcomes the frictional resistance between
the clamped friction plates plus the pre-stressing force in the die springs
(phase 1 in Fig. 2). Then the sliding plates start to move and the die
springs start to compact (phase 2 in Fig. 2). When the die springs are
fully compressed, the end of phase 2, the restoring force preserved in
the compressed disc springs can change the direction of the sliding plate
and provide the required re-centring force for the joint (phase 3 in
Fig. 2). After this point, the sliding plate starts moving towards its initial position and at the end of unloading (phase 4 in Fig. 2), the connection returns to its original configuration.
Note that this joint can be configured to be fully or partially recentring depending on the frictional resistance of the sliding plates,
stiffness of the die springs and the pre-stressing ratio of the system.
Fig. 2 shows the ideal load-deformation response where there is no
residual displacement (and also Fresidual = 0). For the connector shown
in Fig. 1, the specifications can be found using the following equations:
Fslip = 2(nb Fb µ f + Fpr )
(1)
Fmax , loading = Fslip + 2
(2)
max ks
Fmax , unloading = Fmax, loading
Fresidual = Fmax , unloading
4nb Fb µ f
2
max ks
(3)
(4)
The self-centring ratio (rs) of the connector can be calculated using
Eq. (5):
rs =
Fmax, unloading
2
max ks
(5)
A self-centring ratio larger than 1.0 shows a fully self-centring behaviour while a self-centring ratio less than 1.0 shows that the system is
partially self-centring and has some residual deflection. For instance,
Fig. 3 shows the two extreme possible situations. Fig. 3(a) shows a
system with a self-centring ratio more than 1.0 where there is no residual displacement but Fresidual > 0. Fig. 3(b) refers to a system with a
self-centring ratio of 0.0 which is not self-centring and the residual
force is negative. Both scenarios and scenarios in between are possible
depending on the pre-stressing force in the clamping bolts and the
characteristics of the die springs.
Note that to achieve a self-centring behaviour (with the self-centring
ratio equal or larger than 1.0), it is important that the stiffness of the die
springs is aligned with frictional resistance of the plates following Eqs.
(1)–(5). This is also necessary to avoid the ratcheting phenomena in the
rocking wall (when the panel climb while rocking).
2. The concept of re-centring friction hold-downs
3. Joint component tests of the self-centring hold-down
Fig. 1 shows the schematic view of the proposed re-centring holddown connector. This new joint incorporates linear die springs with
sliding slotted friction plates to create a re-centring device. The holddown is comprised of two cover plates attached to the foundation, a
centre plate attached to the wall and friction bolts (accompanied by
disc springs) to sandwich the three plates together. The disc springs are
employed to maintain the clamping force on the bolts thus maintaining
the frictional resistance between the sliding plates. This method has
already been proven to be an efficient way to maintain the friction force
[22].
The two cover plates are used to develop a symmetric friction mechanism where the clamping bolts are under pure tension and no shear
demand is introduced to them. The readers are referred to [23] for more
information about the differences between symmetric and asymmetric
friction connections. The centre plate is wider than the cover plates so
In order to investigate the behaviour of the proposed self-centring
connections on the component level, a series of joint component tests
were conducted. For the tests, a specimen was manufactured with two
cap plates made of mild steel grade 350 MPa and one centre slotted
plate made with Bisalloy 400 (Fy = 1200 MPa). This difference in grade
was chosen based on the findings of Loo et al. [22] to achieve a stable
behaviour in the friction joint. The device had maximum displacement
capacities of 70 mm and 20 mm in tension and compression, respectively.
Fig. 4 shows the general arrangement of the test setup and test
specimen. For this test, a 300 kN MTS machine in the University of
Auckland structure test hall was used.
First, the friction connection without the die springs were tested to
identify the coefficient of friction between the two surfaces. The load
protocol used was three full cycles at the amplitudes of 20%, 40%, 60%,
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A. Hashemi and P. Quenneville
Rocking
Wall
Slotted centre plate
Friction bolts
With disc springs
Stiffeners
Resilient hold-down
Die spring
Rod
Foundation
Fig. 1. The concept of self-centring friction hold-downs.
Fmax,loading
Force
2
figure, the flag-shaped load-deformation curves exhibits a self-centring
behaviour with minimum residual displacement. Additionally, the observed stable and repeatable hysteretic curves confirms the low damage
characteristic of the system given the device maintained its stiffness and
strength over several loading/unloading cycles.
3
Fslip
Fmax,unloading
1
4
Fresidual
max
4. The analytical model for rocking walls with re-centring holddowns
Displacement
The mechanics of the concept of rocking walls with re-centring
hold-downs is shown in Fig. 6. As can be seen, if the wall toes at the
bottom corners are considered as the rocking pivots, the displacement
in each hold-down is related to its lever arm (the distance between the
L
centre of the device and the rocking toe). In other words, 1 = L1 in
2
2
which 1 and 2 are the vertical displacements in the hold-downs. Note
that the elastic deflection of the wall is negligible when compared to the
deflection of the wall due to the rocking motion so it was not considered
in the calculations.
Respecting Fig. 12, the slip threshold of the system Fsys,slip (the force
applied at the top of the wall that triggers the rocking movement) can
be specified using the following equation:
Fig. 2. The ideal load-deformation response of the proposed connector.
80% and 100% of the maximum displacement (30 mm for this specimen) [22]. Fig. 5(a) displays the load protocol and Fig. 5(b) shows the
results of test without implementing the die springs. The clamping bolts
were pre-stressed up to 10 kN each by implementing disc springs under
the bolt shanks and measuring the deflection in the disc springs. The
results showed a stable hysteretic performance with minimum variation
among the load-deformation curves and a coefficient of friction of
µ f = 0.4 between the sliding surfaces.
After this test, the die springs were added to the specimen and each
pre-stressed to 16 kN. Note that before the test, each of the die springs
were separately tested and it was observed that they have a maximum
deformation capacity of 35 mm and a maximum load capacity of 26 kN
which means the stiffness of each die spring is 0.73 kN/mm.
Fig. 5(c) shows the performance of the device. It can be seen that the
joint performed as predicted in accordance with the hysteresis specified
by the proposed capacity equations (Eqs. (1)–(5)). As can be seen in the
Force (kN)
Fsys, slip =
200
100
Force (kN)
300
200
Displacement (mm)
Displacement (mm)
max
max
(a)
(6)
In which h is the height of the wall, W is the gravity loads (including
the self-weight of the wall), LW is the lever arm for the gravity loads and
F1,slip is the slip force of the ascending hold-down (the hold-down in
tension; the left one in Fig. 6). Note that for the descending hold-down
(the hold-down in compression; the right one for the in Fig. 6), the slip
force (F2,slip) is different considering the direction of the applied load.
300
200
1
[WLW + F1, slip L1 + F2, slip L 2 )]
h
-100
(b)
Fig. 3. The extreme load-deformation responses for the proposed connector.
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
Fig. 4. Joint component testing of the re-centring friction hold-down connector: (a) general arrangement of the test setup (b) test specimen.
Fig. 5. Joint component test results of the re-centring friction hold-down connector: (a) the applied load protocol (b) the test result without implementing the die
springs (c) the test results after adding the die springs.
When considering the devices in compression, the frictional resistance
should be higher than the pre-stressing force in the die springs
2nb Fb µf > Fpre to end up with a positive Fslip.
It should be noted that after the devices are activated, the force in
the hold-downs are different and related to the vertical deflection induced in them. Thus, the relative force at the top of the wall at any
stage of rocking can be determined using Eq. (7)
Fsys =
1
(WLW + F1 L1 + F2 L2)
h
(7)
F1 and F2 are hold-down forces for the left and right devices for the
left to right direction of rocking shown in Fig. 12. The accuracy of the
proposed equations is assessed in the next section by comparing the
analytically calculated values with the experimentally obtained data.
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
b
schematic view of the wall and a picture from the test setup. This
section describes the components of the test rig shown in Fig. 6(b) and
the instrumentation details.
Fsys
5.1.1. The CLT rocking wall panel
A five-ply CLT panel was considered for testing. The wall is 6 m
high, 2025 mm wide and 200 mm thick. It has three longitudinal layers
oriented along the height and two transverse layers. Note that the
thickness of the planks used was 40 mm. The grade of all timber planks
was MSG 8 [24] grade with a density of 510 kg/m3 resulting in a selfweight of 1.2 tonnes.
Note that in this concept, the ductility is provided by the re-centring
hold-downs and the rest of the structure is meant to remain elastic.
Previous studies have demonstrated that long screws can provide relatively stiff and yet economical connections [25,26]. Accordingly, long
self-tapping screws were considered to attach a steel bracket to the
notch implemented at the bottom corners of the panel and then the
hold-downs were pinned to that steel bracket. The use of long selftapping screws has previously been proven to have the sufficient stiffness in CLT applications [27,28]. Fig. 8 displays the cross section of the
CLT panel and the details of the top steel bracket and the screws used in
the connection. The hold-downs were connected to the foundation
plates using M20 bolts (grade 8.8) resulting in a double-pin design for
the device to minimize any possible eccentricity and the resulting
bending moments. During the installation of the hold-downs, a 10 mm
gap was considered between the devices and the foundation plates to
allow for a minor displacement in compression.
LW
h
W
F1
F2
L1
L2
Fig. 6. Rocking wall with re-centring hold-downs.
5. Large scale testing of a rocking Cross Laminated timber (CLT)
wall with re-centring hold-downs
5.1. Structural components
5.1.2. The hold-downs and the foundation plates
Two devices similar to the one shown in Fig. 4(b) were considered
Fig. 7 shows the general arrangement of the test setup including a
Out of plane
restraint
Actuator
Load transfer
channels
Re-centring
hold-down
Shear key
Foundation
plates
Fig. 7. Experimental testing of a rocking CLT wall with re-centring hold-downs.
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
10 mm thick stiffeners
for out of plane bending
reinforcement
20 mm tight fit hole
50
200 100
20
35
12
34
12
13
100
55
10
157.5
157.5
50
315
60
60
60
75
60
315
(a)
(b)
Eight Ø=11 mm
self-tapping screws
with 550 mm
Top steel
bracket
Pin
M20 bolt
590
Top steel
bracket
(d)
345
(c)
Fig. 8. The connection between the specimen and the CLT panel (all dimensions are in mm): (a) the top steel bracket (top view) (b) the top steel bracket (front view)
(c) the connection between the top steel bracket and the panel (d) the hold-down installed in place.
as the hold-down connectors for the CLT panel. As mentioned, a double
pinned assembly was used to avoid any bending demand on the holddown. Nevertheless, the devices could be designed similar to a conventional steel connector if there was any bending demand. It should be
noted that the pin and bolt assemblies at the top and the bottom of the
device have a tolerance of 0.02 mm between the pin/bolt and the pin
hole to minimize any possible slack in the system during the tests. Note
that the bottom steel brackets were bolted to the plates that were already welded to the foundation plates. The 20 mm diameter pins were
made of high strength steel grade Fy = 690 MPa.
As mentioned earlier, the hold-downs could provide a maximum
deflection of 70 mm in tension and 20 mm in compression. This is
because when the panel rocks in one direction and one hold-down ascends, the deflection in the opposite descending hold-down is relatively
smaller (which is related to the lower applicable eccentricity).
For the foundation, three 36 mm thick steel plates (made with mild
steel grade 350 MPa) were considered under the panel. Each of the
foundation plates was attached to the concrete strong floor using posttensioned rods at the corners. The post-tensioning load used for each
rod was 120 kN. With this configuration, the frictional resistance between the steel plates and the concrete floor beneath avoided any potential horizontal sliding during the tests. Fully threaded holes were
also considered in the central plate to attach the shear key. Fig. 9 shows
a schematic view of the foundation plates. Note that to provide a similar
bearing condition for both assumed pivoting rocking points, a piece of
timber was added between the two right foundation plates (see Fig. 9)
during the tests.
5.1.3. Shear keys
Previous experimental studies on CLT platform structures have demonstrated that the sliding movement between the wall panel and the
foundation or the floor below has the largest contribution in the lateral
drift of the structure [7,8,29]. This sliding, which is not recoverable,
causes slackness in the system making it significantly vulnerable to the
aftershocks and even the system may not be able to tolerate long
duration seismic events. Therefore, a mechanism which can transfer the
lateral shear forces (with minimum sliding movement) while accommodating the uplift caused by the rocking motion of the wall is
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
Eight M20 threaded holes for
connecting the shear keys
2810
860
810
860
500
120
810
120
500
200
3@100
140
140
32 mm holes to anchor the plates
to the strong floor by 27 mm
diameter posttensioned rods
Position of the
CLT panel
Plate welded to attach the
hold-downs
Fig. 9. The schematic view of the foundation plates (dimensions in mm).
necessary.
There were some attempts before to provide such shear transferring
mechanism such as the use of steel angles at the rocking toes [30] or
providing a rod-to-plate contact shear key [14]. In this paper, a shear
transferring device is proposed to compensate for the issue mentioned.
This device works similar to the conventional bolted connectors but the
bolt holes are profiled in a way that can de-couple the vertical movement of the wall panel while transferring the shear forces at all wall
rocking positions. In other words, when the wall rocks, the bolts are
always in contact with the slanted surfaces in the steel plate.
The geometry of the bolt holes can be determined based on the
geometrical properties (aspect ratio) of the rocking panel. For this test,
the holes were profiled in a way that can accommodate 5% of lateral
drift during the rocking movement of the wall. Two shear keys were
used for this experiment (one at each side of the panel) in a way that
four bolts (M20 grade 8.8) were implemented in the wall and were in
contact with the slanted surfaces of the two plates with profiled holes. A
commercial grease was used between the contact surfaces of the shear
key and the panel to avoid any frictional resistance between them.
Fig. 10 shows the shear key used.
5.2. Test equipment and instrumentation
A 300 kN hydraulic jack was used to apply the lateral loads to the
wall panel. The lateral load from the actuator was applied to the wall
using two back-to- back C-channels bolted to the CLT wall. 50 kN disc
springs were used under the bolt shanks to develop sufficient frictional
resistance between the load transfer channels and the CLT panel. The
bolts were tightened until the disc springs were flattened. Considering
the coefficient of friction between steel and timber (approximately a
minimum of 0.4), the resistance was almost four times larger than the
maximum anticipated lateral force coming from the actuator.
Fig. 11 (a) shows the load transfer channels and the lateral support
employed. To minimize the potential out-of-plane movement of the
wall during the in-plane testing and keep the wall in the vertical position, two pieces of timber (1200 mm by 140 mm, approximately) were
bolted to the lateral supports (see Fig. 11(a)).
Fig. 11(b) illustrates the location of the instruments used to measure
the force and displacements for the tests. A 300 kN load cell was
mounted to the hydraulic jack to measure the force. Linear Variable
Displacement Transducers (LVDTs) were implemented in different locations to measure the deflections. As can be seen in the figure, one
LVDT was attached to each of hold-down to measure the vertical displacement in the devices. Furthermore, one LVDT was connected to
each of the shear keys to measure the relative movement between the
wall and the foundations plate. In addition, a larger LVD capable of
recording displacements up to 150 mm was employed to measure the
main horizontal displacement of the CLT wall at the actuator level.
Finally, two displacement gauges were placed at the top steel
brackets to detect the potential deformation of the screwed connections
and to monitor that if this configuration was stiff enough.
500
10
90
100
100
100
90
10
10.5 13
200 188
80
21
12
Weld
(a)
6. Testing configurations
Different tests were performed on the CLT rocking wall with recentring hold-downs. In order to investigate the performance of the
system in different situations, the pre-stressing on the clamping friction
bolts was kept constant during the tests while the pre-stressing force
applied to the die springs was changed. As mentioned in the previous
section, the force in the clamping bolts was tuned by measuring the
deflection in the disc springs under the bolt shanks using a digital micrometre. These disc springs had 40 kN maximum load capacity and
had a linear load-deformation behaviour.
(b)
Fig. 10. The shear key (dimensions in mm): (a) geometry (d) implementation.
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
(a)
Load cell
Main LVDT
Shear key LVDT
(b)
Lateral supports
Bolts with disc
springs
Load transfer
channels
Hold-down
LVDT
3315 mm
(c)
2025 mm
Fig. 11. Test equipment and instrumentation: (a) the wall panel (b) shear key instrument (c) hold-down instrument (on the back).
1 mm/s respecting the test equipment and limitations of the hydraulic
jack used.
Table 1
Test matrix.
Test code
Pre-stressing force in
each friction bolt (kN)
Pre-stressing force in
the die springs (kN)
Lateral deflection of
the wall (mm)
Test
Test
Test
Test
10
10
10
10
4
5
7
9
100
65
58
40
1
2
3
4
7. Experimental results and discussions
7.1. Cyclic performance of the system
Altogether, four tests were performed on the specimen in which the
related cyclic load-deformation responses are shown in Fig. 13. In the
graphs shown, the readings from the load cell (mounted on the hydraulic jack) are plotted against the displacements measured by the
main LVDT (at the actuator level).
A bi-linear load-slip behaviour is observable for all test which shows
the potential for having a self-centring system. Note that in Fig. 13,
some residual displacements are observable. However, this residual
displacement could have avoided if stiffer die springs were used for the
dampers. Note that to achieve a fully self-centring behaviour, the prestressing force applied to the die springs needs to be increased, however, then the leftover displacement capacity in the springs would not
be sufficient to provide the desired uplift in the hold-downs. Therefore,
stiffer die springs that can have more pre-stressing force (while having
the same displacement capacity) could be a solution in this case.
It can be seen that the behaviour of the system is stable over multiple
cycles of loading and unloading with insignificant variability in the
curves. Note that for this test, the only contributed gravity load was the
self-weight of the wall panel plus half of the weight of the actuator and
the weight of the load-transfer channels (overall 12.5 kN). Therefore, it is
demonstrated that by using the proposed system, there is a potential for
having a self-centring system is without applying extra dead loads (or
any extra vertical resistance such as post-tensioned cables). Also, the
symmetrical hysteretic performance of the wall panel confirmed the
accuracy of the method used to tighten the clamping bolts.
Displacement (mm)
Table 1 provides the test matrix of the conducted test. A displacement-control approach was chosen for each test where the target displacement was applied at the actuator level by adjusting the pressure in
the hydraulic jack and reading the deflection from the main LVDT.
Fig. 12 shows the applied load schedule that included three fully
reversed cycles of loading and unloading applied at amplitudes of 15%,
45%, 70% and 100% percent of the target lateral displacement. Different target displacements were considered given the limitations of the
hold-downs and the testing equipment. Fig. 11 displays the applied
displacement history for Test 1. The loading rate was approximately
100
75
50
25
0
-25 0
-50
-75
-100
400
800
1200
1600
2000
2400
2800
Time (sec)
Fig. 12. The applied displacement schedule for Test 1.
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A. Hashemi and P. Quenneville
Fig. 13. Cyclic performance of the system: (a) test 1 (b) test 2 (c) test 3 (d) test 4.
It should be pointed out that the maximum displacement at the top
was different for the four tests because of the different pre-stressing
force applied to the die springs used and their limited leftover maximum displacement. In other words, the limitation for the proposed recentring hold-down in this research could be the force and deflection
capacity of the die springs. When high Fslip for the hold-downs are required, which means high pre-stressing in the die springs, the deflection
limit of the hold-down could be a restraint. One solution to address this
issue is to increase the number of die springs (and rods) working in
parallel. However, this would increase the size and manufacturing cost
of the hold-down. More to the point, after applying the desired prestressing force to the die springs, there must be enough displacement
capacity remaining in the springs to provide the required deformation.
In sum, the provided solution has some limitations so might be more
suitable for low- to mid-rise structures. Note that the system can be
designed for any desired specifications, but the feasibility of the concept
may need to be investigated for each specific design case.
No damage was observed to the wall itself nor in any other components in the system and also no ratcheting phenomena was noticed
during the tests. This phenomena usually occurs in shear walls with
friction dampers when the vertical loads are not sufficient to bring back
the sliding plates to their original position and as a result, the wall panel
rises on the dampers [14]. In the current study, the restoring force
preserved in the semi-compacted die springs was sufficient to bring
back the moving plates to their initial position and avoid this phenomena.
When considering the largest applied lateral load at the actuator
level (around 30 kN in Fig. 13(a)), the elastic bending and shear deformation of the CLT panel at that height is less than 0.4 mm which is
negligible when compared to the lateral deflection of the wall due to
the expansion of the hold-downs. This fact confirms the assumption
made to consider the wall and other components as rigid bodies when
developing the capacity prediction equations. Note that at the maximum of the horizontal recorded deflections, a small vertical force is
developed because of the slight rotation of the swivel bearing at the
actuator level (see Fig. 6). This load, that is relatively small in magnitude, does not have an influence on the overall behaviour of the system
and can be neglected.
Fig. 13(d) shows a comparison between the experimental data and
the analytically calculated load-deformation response from the proposed equations in the last section. It can be seen that the analytical
prediction matches the experimental results well despite some minor
differences. Table 2 shows the geometric characteristics of the test
specimen, the target load-deformation relationship for the hold-downs
and the expected hysteretic behaviour of the wall panel. It should be
emphasized that similar to the hold-down component tests, the target
friction force in the hold-down was achieved by tightening the bolts
while measuring the deflection of the disc springs (using a digital micrometre) up to the target force in the bolts.
In Summary, the experimental test results demonstrated a stable
cyclic performance with a self-centring tendency. Therefore, it can be
concluded that the system has the potential to be considered as a low
damage structural system given the observed stability in the hysteretic
curves.
7.2. Performance of the hold-downs and shear keys
Fig. 14 shows the hold-downs in tension (the left hold-down) and
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
Table 2
Parameters for Test 4 shown in Fig. 13(d).
Geomtric chracteristics of the system
Target hold-down specificaitons
Target load-deformation behaviour of the system
Item
Value
h (height of the wall, mm)
b (the width of the wall, mm)
L1 (lever arm for the ascending hold-down, mm)
L2 (lever arm for the descending hold-down, mm)
LW (lever arm for the gravity loads, mm)
F1,slip (slip force in the ascending hold-down, kN)
F1,loading (hold-down force in loading, kN)
F1,unloading (hold-down force in unloading, kN)
Fsys,slip (horizontal slip force of the wall, kN)
Fsys,loading (horizontal force in loading, kN)
Fsys,unloading (horizontal force in unloading, kN)
sys (deflection of the wall at the actuator level, mm)
3315
2025
1508
173
668
25
36
5
10
29
16
40
Fig. 14. deformed shape of the hold-downs in tension and compression.
compression (the right hold-down). As can be seen in the figure, the
displacement demand for the hold-down in tension is much larger than
the one in compression. The double-pin design for devices worked as
anticipated and no significant relative lateral movement between the
sliding plates was observed.
Fig. 15(a) displays the vertical displacements recorded by the holddown LVDTs in Test 1 (a similar trend was recorded for all tests) corresponding to the left and right hold-downs. From the figure, it is apparent that the hold-downs performed as expected respecting the load
schedule applied to the wall. In other words, the curves confirmed the
fact that there was no damper ratcheting effect in the system. This type
of behaviour can be attributed to the effect of die springs.
Furthermore, the low damage characteristic of the system was
confirmed when considering the large cumulative travel distance for
each hold down (more than 20 m including the hold-down component
tests). Note that the joints were opened up and inspected after the tests
and no significant damage was observed in any component.
The values recorded by the gauges measuring the displacement
between the top hold-down steel bracket and the wall were less than
0.3 mm for all tests. This shows that the screwed connection detail
chosen to connect the bracket to the CLT panel was efficient. Note that
for low damage concepts, all of the components in the system (other
than the weak links or in this case, the hold-downs) have to remain
elastic up to the design level demands magnified by an appropriate
over-strength factor. In case of the presented concept, the screwed
connection is part of the capacity-designed components. The test results
showed the efficiency of the design given the very low recorded deformations between the CLT panel and the hold-down top bracket.
The deformed shape of the shear key is shown if Fig. 14 and the
sliding movement between the shear keys and the foundation plates are
illustrated in Fig. 15(b). On the basis of the data recorded, it can be seen
that the horizontal displacement of the wall at the base is less than
2 mm while the wall was able to achieve up to 3% lateral displacement
at top (100 mm). This means that an almost pure rocking deflection
mode was achieved using the proposed shear key. In other words, the
shear keys efficiently avoided any sliding at the wall base while accommodating the uplift caused by the rocking movement. No damage
to the bolts or the bracket of the shear keys was noticed after the tests.
This also confirms the low damage characteristic of this shear-transferring device.
Overall, the experimentally obtained data would seem to suggest
that the proposed structural system has the merit to be potentially
considered as an effective low damage solution for timber construction.
8. Conclusions
This paper presented the experimental test results of a rocking Cross
Laminated Timber (CLT) wall with a new friction-based energy dissipation device used as hold-down connectors. The integration of die
springs with a conventional friction joint added the advantage of recentring to the concept. Based on the concept introduced, a prototype
device was manufactured and subjected to cyclic displacement-control
testing. Furthermore, large-scale cyclic tests were performed on a
rocking CLT wall with re-centring hold-downs. The test data demonstrated repeatable hysteretic performance without any stiffness degradation that confirmed the low damage characteristic of the concept.
In the tests, the hold-down worked as the fuses to provide the required
geometrically non-linear behaviour and the rest of the system remained
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Engineering Structures 206 (2020) 110166
A. Hashemi and P. Quenneville
Displacement (mm)
50
Left hold-down
Right hold-down
40
30
20
10
0
-10
0
200
400
600
-20
800
1000
1200
1400
1000
1200
1400
Time (sec)
(a)
1.5
Displacement (mm)
1
0.5
0
-0.5 0
200
400
600
800
-1
-1.5
-2
Shear key 1
Shear key 2
-2.5
Time (sec)
(b)
Fig. 15. Cyclic performance of the components: (a) hold-downs (b) shear keys.
elastic. The novel type of shear keys adopted performed as predicted
allowing the wall panel to move in the intended manner while the shear
forces being efficiently transferred to the foundation. Overall, the results of this study showed that the presented concept has the potential
for application in low damage CLT structures. For implementation of
the devices, each one can be tuned and tested before installation to
make sure that they perform to the design specifications. Nevertheless,
the long-term reliability of the system may need to be studied in future
in to make sure that the design parameters stay reasonably close the
initial design values. Furthermore, the performance of the proposed
concept on the system level is planned to be investigated n future by
developing case-studies and performing non-linear dynamic time-history simulations. This is also necessary make sure that the forces flow as
intended the CLT walls and then the devices.
Future research will also include dynamic testing to assess the behaviour of the system at higher load rates. Furthermore, the use of
different types of die springs with different elongation capacities can be
considered for future testing to achieve a fully self-centring behaviour.
Acknowledgement
The authors would like to thank the Earthquake Commission
Research Foundation (EQC) of New Zealand for the financial support of
the presented research under the grant number “EQC 15/U710”.
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CRediT authorship contribution statement
Ashkan Hashemi: Conceptualization, Methodology, Investigation,
Validation, Formal analysis, Writing - original draft, Writing - review &
editing, Project administration. Pierre Quenneville: Supervision,
Resources, Project administration, Funding acquisition, Writing - review & editing.
Declaration of Competing Interest
We declare that we have no conflict of interest.
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