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Bending and shear performance of Australian Radiata pine cross-laminated timber
Construction and Building Materials 232 (2020) 117215
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Bending and shear performance of Australian Radiata pine cross-laminated timber S. Navaratnam a,⇑, P.B. Christopher b, T. Ngo b,⇑, T.V. Le b a b
School of Engineering, RMIT University, Building 10| Level 13| Room 17A, 376-392 Swanston Street, Melbourne VIC 3001, Australia Department of Infrastructure Engineering, The University of Melbourne, Australia
h i g h l i g h t s Bending and shear behavior of Australian radiata pine cross-laminated timber. FEM analysis to predict bending and shear strength of CLT. Failure modes of CLT under out-of-plane bending and shear. Comparison of bending and shear stiffness of CLT between tests and theoretical.
a r t i c l e
i n f o
Article history: Received 15 April 2019 Received in revised form 7 October 2019 Accepted 10 October 2019
Keywords: Cross-laminated timber Bending strength Shear strength Failure mode Finite element model Radiata pine
a b s t r a c t Cross Laminated Timber (CLT) is increasingly being used in commercial and residential construction in Australia due to its inherent strength and sustainability credentials. Until recently, infrastructure building projects using CLT have been reliant on imported products from overseas manufacturers. There is now a viable Australian grown and fabricated CLT product from Radiata pine. This paper summarises the experimental results on the mechanical behaviour of Australian Radiata pine CLT panels in out-of-plane bending and shear. Three-layer 105 mm thick panels and five-layer 145 mm panels with three different spans were tested. The results demonstrate strong correlations to existing theoretical models and were used to validate the finite element model (FEM) developed in this research. The experimental results showed that the average bending stiffness of the CLT panels were marginally greater than theoretical values. The maximum bending strength for Radiata pine exceeded the characteristic strength of 14.0 MPa for grade XLG1 external laminas, with three-layer CLT samples averaging 28.7 MPa and five-layer CLT samples averaging 26.8 MPa. The maximum observed shear stresses ranged from 1.55 MPa to 2.18 MPa, which also exceeded the rolling shear characteristic strength of 1.2 MPa for the feedstock. The results also highlighted that the shear strength decreases with an increasing thickness of the CLT panel. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction The construction industry demands continual improvements for building materials that are fast and safe to erect, cost effective and sustainable [1–3]. CLT shows potential to address many of these requirements in the industry, with a high degree of prefabrication, high in-plane and out-plane strength and stiffness, and good acoustic and thermal performance [4,5]. CLT is an engineered wood panel utilising at least three layers of timber boards glued
⇑ Corresponding authors. E-mail addresses: [email protected] (S. Navaratnam), [email protected]. au (P.B. Christopher), [email protected] (T. Ngo), [email protected]. au (T.V. Le). https://doi.org/10.1016/j.conbuildmat.2019.117215 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.
together in orthogonal directions. This combination of layers eliminates much of the natural variability and the influence of defects in timber products when compared to plain sawn timber. CLT panels are frequently fabricated using timber offcuts or lower grade feedstock from other structural timber productions, thereby having a lower embodied energy, greenhouse gas emissions and high carbon sequestration compared to typical concrete construction [6]. As a result of these advantages, CLT is being used more widely in building construction [4,7,8]. CLT panels are made from different timber species that depend on local resources such as Kiri, Katsura, Sugi, Hinoki, Buna spruce pine (Europe and Canada) and Radiata pine (Australia and New Zealand). Previous research has focused on the flexural and shear performance of a variety of timber feedstock for CLT panels
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
[4,9,10–16]. Hindman and Bouldin [9] performed four-point bending tests in order to evaluate the bending and shear properties of Southern pine CLT. Their results indicated an equal or higher shear resistance compared to the standard specified values of shear (i.e. GAeff = 23 MN/m) and bending (i.e. EIeff = 9.7 MNm) resistance [17]. Similarly, four-point bending tests were used to evaluate the stiffness and strength properties in the perpendicular and parallel directions to the principal plane of the CLT panels, which were made using spruce boards [11,14]. Sikora et al. [14] assessed the effect of the panel thickness of Irish Sitka spruce CLT and found that the bending and rolling shear strength decreased with an increasing panel thickness. The bending creep performance of engineered wood products has been assessed by Park, et al. [10], where CLT panels from five different timber species were tested and they found that the creep deformation perpendicular to the grain was decreased by cross laminating. Fortune and Quenneville [12], Lewis et al. [18], Woodco [19], Iqbal [20] all investigated the shear and bending performance of CLT panels fabricated from New Zealand Radiata pine. These previous studies found that the bending and shear strength, and stiffness of CLT panels were equal to or higher than the standard specifications. However, there is limited research that has assessed the bending and shear performance of CLT panels fabricated using Australian Radiata pine. This research investigates the bending and shear performance of Australian Radiata pine CLT panels using four-point bending tests and finite element (FE) analysis. Three different spans with two different thicknesses of CLT panels were tested to evaluate the bending and shear strength, and stiffness through experimental tests. The experimental results of this study also used to validate the FEM and predict the shear and bending performance of CLT panels with seven-layer.
2. Experimental tests and materials 2.1. Specimen and material details Experimental tests were performed to evaluate the strength and stiffness of CLT panels for both three and five-layer CLT build ups. The CLT panels were manufactured by XLam in 2018 with the three-layer 105 mm thick panel (CL3/105) made up with 35/35/35 mm layups, whilst the five-layer 145 mm thick panel (CL5/145) is made up of 35/20/35/20/35 mm layups. Each specimen was manufactured from 35 90 mm and 20 140 mm Australian Radiata pine timber and bonded with Purbond polyurethane adhesive without edge bonding. The thickness, width, modulus of elasticity (MOE), shear modulus and the lamination grades of the CLT test specimens are detailed in Table 1. Both the top and bottom outer layers are manufactured with XLG1 (comparable to F7) structural grade Radiata pine timber, while the inner layers are made with XLG2 (comparable to F5) structural grade Radiata pine timber [21]. The specified average MOE for the outer layers and inner layer is 8 GPa and 6 GPa, respectively. The average density of the CL3/105 panels is 492 kg/m3 with a coeffi-
cient of variation (COV) of 10%. The average density of the CL5/145 panels was 483 kg/m3 with a COV of 11%. 2.2. Test methods The CLT panels were tested using an MTS hydraulic actuator with a capacity of 500 kN and the load configurations were in accordance with I.S. EN 16351[22]. To derive the strength and stiffness properties, the standard I.S. EN 16,351 [22] was used, which specifies the span and number of test specimens based on the thickness (h) and width of the specimen. Four-point bending was applied over a span of 20 h and 28 h to derive the bending strength and stiffness properties, whereas 12 h spans were used to determine the shear (rolling) strength and stiffness perpendicular to the plane. Table 2 details dimensions of the specimens and the purpose of each test. The specimens were stored at 20 ± 2 °C and ~65% relative humidity for a minimum of 72 h prior to the experiments. The moisture content of each timber layer was measured and found to be 10.8% on average with a COV of 12%, which is above the manufacturer’s specification of not less than 8%. As specified by I.S. EN 16,351 [22] a constant displacement-controlled loading rate was applied and adjusted for each span and thickness of specimen to ensure the maximum load was achieved within 300 ± 120 s. Three laser transducers (LS1, LS2 and LS3) were used to measure and calculate the local and global displacements as shown in Fig. 1. The laser transducers LS1 and LS3 were positioned such that the central gauge length was five times that of the depth of the CLT panel. The global displacement was measured by the laser transducer LS2 at the mid-span of the test specimen. The local displacement was determined by deducting the average displacement obtained from LS1 and LS3 from the global displacement (LS2). Multiple cameras were used to assess the lateral and vertical movement of the panels, and capture the crack propagation through the thickness. The cameras were positioned near the supports, shear zones and the shear-free zone (Fig. 1). Digital image correlation (DIC) analysis was performed to assess the out-ofplane movement panel, in-plane strain and crack growth of CLT [23]. Based on this analysis, its noted that the lateral and vertical movement at the supports was negligible until 90% of the ultimate load was reached. Specifically, a sudden lateral displacement between 10 and 15 mm occurred at the supports during the shear test when the ultimate load was reached. This movement does not affect the global and local stiffness of the panel as the flexural rigidities were calculated between 10% and 40% of the ultimate load using Eqs. (1) and (2). 3. Results and analysis 3.1. Bending test results The global load-displacement curve for the ten CL3/105 samples tested over a span of 2940 mm is shown in Fig. 2. It can be
Table 1 CLT test specimen details and material properties. Specimen details Type
Lamination grade
Lamination thickness (mm)
Width (mm)
Average density (kg/m3) /COV (%)
CL3/105 CL5/145 Mechanical properties Lamination grade MOE (MPa)
XLG1/ XLG2/ XLG1 XLG1/ XLG2/ XLG1/ XLG2/ XLG1
35/35/35 35/20/35/20/35
520 520
492/10 483/11
Parallel to the grain Perpendicular to the grain Longitudinal (G0) Rolling shear (GRS)
XLG2 6000 200 400 40
XLG1 8000 267 533 53
Shear modulus (MPa)
3
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215 Table 2 CLT panel test specimens loaded perpendicular to the plane. Specimen label
Number of samples
Number of layers
Thickness (mm)
Span (mm)
Loading direction
Determined properties
CL3/105/1260 CL5/145/1740 CL3/105/2100 CL3/105/2940 CL5/145/2900 CL5/145/4020
9 10 10 10 10 10
3 5 3 3 5 5
105 145 105 105 145 145
1260 1740 2100 2940 2900 4020
Perpendicular Perpendicular Perpendicular Perpendicular Perpendicular Perpendicular
Shear strength & stiffness Shear strength & stiffness Bending strength & stiffness Bending strength & stiffness Bending strength & stiffness Bending strength & stiffness
to to to to to to
plane plane plane plane plane plane
Fig. 1. Testing set-up: (a) Experimental test set-up; (b) Schematic diagram and nomenclature.
observed that the ultimate load (Fmax) of each test specimen varies between 46.7 kN and 56.7 kN with a COV of 7%. Fig. 2 also shows that all samples exhibited brittle failure at the ultimate load. The average global and local force–displacement curves for each bending test sample (i.e. CL3/105/2100, CL3/105/2940, CL5/145/2900, CL5/145/4020) were obtained from the tests and are shown in Fig. 3. It can be observed that the average ultimate load (Fmax) for the five-layer CLT panel was about 10% higher than the threelayer CLT panel with the same span to thickness ratio. This indicates that increasing the thickness of the panel whilst maintaining the same span-to-thickness ratio does not improve its flexural capacity. Similar phenomenon was in agreement with the finding made by Sikora, et al. [14] for the CLT panels made from Irish Sitka spruce. The most common failure modes observed in the bending tests were tensile failure of the lower lamina, delamination between the lower laminas and tensile failure at the finger joints, as shown in Fig. 4. The global displacement which accounts for both shear and bending deformations was used together with the force for each sample to derive the global stiffness (EImg). In a similar way the local force–displacement relationship was used to calculate the local flexural stiffness (i.e. bending stiffness in the absence of shear deformation) in the absence of shear stresses (EIml). Based on EN
408 [24], the following equations were used to derive the global and local stiffness, which were similar to those used by Sikora et al. [14] and He et al. [4]: 2
EIml ¼
al1 ðF 2 F 1 Þ 16ðw2 w1 Þ
EImg ¼
3aL2 4a2 3a 1 48ðwF 22 w 5Gbh Þ F 1
ð1Þ
ð2Þ
where: a is the distance between the loading head and the nearest support; l1 (5 h) is equal to the gauge length for the local displacement measurement; L is the span (Fig. 1); F1 and F2 are the loads corresponding to 10% and 40% of the ultimate load (Fmax); w1 and w2 are the displacements corresponding to loads F1 and F2, respectively; G is the shear modulus; and b is the width of the panel (520 mm). Two global stiffnesses were determined by taking the shear modulus as both infinite and 533 N/mm2 (see Tables 3 and 4). The highest stiffness obtained for the three-layer CLT panels were calculated as 5.39E + 11 Nmm2 using the local displacement and 4.67E + 11 Nmm2 using the global displacement, with an infinite shear modulus. Both the local and global stiffnesses were higher in the fivelayer CLT panels, namely about 93% and 119%, respectively. These
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Force (kN)
4
CL3/105/2940_Average CL3/105/2940_T1 CL3/105/2940_T2 CL3/105/2940_T3 CL3/105/2940_T4 CL3/105/2940_T5 CL3/105/2940_T6 CL3/105/2940_T7 CL3/105/2940_T8 CL3/105/2940_T9 CL3/105/2940_T10
60 50 40 30 20
10 0 0
10
20
30
40 50 60 Global displacement (mm)
70
Fig. 2. Global force-displacement response for sample CL3/105/2940.
(a)Global displacement
(b)Local displacement
Fig. 3. Average force-displacement response for bending test samples: (a) global force-displacement; (b) local force-displacement.
results indicate that an increase in the thickness of the CLT panel by 40% will result in doubling of the bending stiffness. The bending strength (fb) of the CLT panels was obtained using the Shear Analogy theory which takes into consideration the shear deformation of the cross laminated layers and is tabulated in Tables 3 and 4. The flexural strength of the outer laminates (fb) of the CLT panel can be derived using Eq. (3).
fb ¼
M max E1 h=2 EIml
ð3Þ
where: Mmax is the maximum bending moment; E1 is the MOE of the outer layer; and EIml is the local stiffness obtained from Eq. (1). The highest fb recorded for the three-layer and five-layer panels were 33.45 MPa and 36.03 MPa, respectively. The average fb for the five-layer CLT panel (i.e. 26.84 MPa) was approximately 15% higher than that of the three-layer panel (i.e. 23.41 MPa) with the same span-to-thickness ratio. This indicates that increasing
the panel thickness does not improve the bending strength of the outermost layer for panels with the same span-to-thickness ratio. Further, the average fb values for the Australian radiata pine CLT are approximately between the bending properties of machine graded pine (MGP) 10 and 12 as specified in Australian Standard AS 1720.1 [25]. 3.2. Experimental and theoretical comparison of the bending stiffness Three methods were used to calculate the theoretical bending stiffness (EIeff) for each CLT panel; the Shear Analogy Theory, Modified Gamma Theory and Composite Theory (k method). Eqs. (4) and (5) as specified by Karacabeyli and Douglas [26] were used to calculate the effective bending stiffness based on the Shear Analogy theory (EIeff, shear) and Composite theory (EIeff,comp), respectively. Eq. (6) was used to determine the effective bending stiffness using the Modified Gamma theory, which does not con-
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
5
Fig. 4. The most common failure modes from the bending tests: a) tension failure; b) delamination; c) tension failure at the finger joints.
sider shear deformation and is founded on the Mechanically Jointed Beams theory EN 1995 [27].
EIeff ;shear ¼
n X i¼1
EIeff ;Comp ¼ Eeff :
EIeff ;Gamma ¼
n X hi Ei :Ai :zi 2 þ 12 i¼1 3
Ei :bi :
b:htot 12
n X
ð4Þ
3
ð5Þ 3
Ei :bi :
i¼1
hi 12
! þ
Xn i
ci :Ei :bi :hi :zi 2
ð6Þ
where: bi is the width of each layer; hi is the thickness of each layer; zi is the distance between the center point of each layer and the neutral axis; Ei is the MOE; htot is the thickness of the panel; and Eeff = 30 k1E, k1 is determined by the load configuration. The connection efficiency factor ci , which can be derived from Equation (7) below, is non-zero only for longitudinal layers and equal to unity for the middle layer.
ci ¼ 1 þ
p2 :Ei :bi :hi
L2 eff : G:bj =hj
!1 ð7Þ
where: Leff is the effective length of the beam; and j refers to the transverse layer connecting the ith layer with the central layer.
Fig. 5 shows a comparison between the theoretical and experimental bending stiffness. The theoretical values of EIeff,shear, EIeff,comp and EIeff,Gamma were found to yield similar estimations of the bending stiffness. The experimentally determined global bending stiffness EImg was consistently less than the experimentally determined local bending stiffness EIml, however both experimental values were found to be higher than any of the theoretical bending stiffnesses. Similar results were reported by Sikora et al. [14] and He et al. [4]. Fig. 5 also shows a comparison between the theoretical and experimental bending stiffness, and illustrates that the EImg with an infinite shear modulus gives less differences than EIml and EImg with a finite shear modulus of 533 MPa. This indicates that the theoretical equations can be accurately predict the global bending stiffness with shear and bending deformation. The average EImg with an infinite shear modulus (4.53E + 11 Nmm2) was about 18% higher than the theoretical bending stiffness in the three-layer CLT panel, whilst in the five-layer CLT panel, it was about 5% higher when compared to the theoretical bending stiffness. This indicates that the theoretical prediction of EImg with an infinite shear modulus is possibly more accurate with an increased number of layers in the CLT panel. 3.3. Shear test results The average global and local force-displacement relationship from the shear tests are shown in Fig. 6. The ultimate load (Fmax)
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Table 3 Bending test results for the tree-layer CLT panels. Specimen details
Test No
Experimental bending stiffness, EI 1011 (Nmm2) EIml
EImg with G =1
EImg with G = 533
CL3/105/2100
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
5.31 4.93 4.94 5.32 4.96 5.10 4.97 5.04 5.39 5.13 5.11 0.17 0.03 5.13 5.03 5.41 5.22 5.06 5.19 5.16 5.00 5.02 5.17 5.14 0.12 0.02
4.50 4.26 4.33 4.56 4.29 4.44 4.33 4.35 4.67 4.42 4.42 0.13 0.03 4.66 4.55 4.78 4.74 4.63 4.77 4.70 4.55 4.60 4.65 4.66 0.08 0.02
4.66 4.40 4.47 4.72 4.43 4.59 4.47 4.50 4.84 4.57 4.57 0.14 0.03 4.75 4.63 4.87 4.84 4.72 4.86 4.79 4.63 4.68 4.74 4.75 0.09 0.02
Mean Std COV CL3/105/2940
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
Mean Std COV
Mmax (kNm)
fb,0 (MPa)
Failure mode
38.21 34.73 24.16 27.58 31.35 40.55 34.52 31.23 31.00 30.13 32.35 4.84 0.15 27.18 31.59 31.58 26.06 26.54 27.66 29.50 29.33 26.58 30.45 28.65 2.11 0.07
30.20 29.56 20.53 21.78 26.52 33.45 29.19 26.02 24.17 24.69 26.61 4.02 0.15 22.27 26.37 24.49 20.97 22.02 22.39 23.99 24.66 22.26 24.73 23.41 1.67 0.07
Tensile & Delamination Tension Tension Tension Tension Tensile & Delamination Tensile & Delamination Tension Tension Tension
Tension Tension Tension Tensile & Delamination Tension Tension Tensile & Delamination Tension Tension Tension
Table 4 Bending test results for the five-layer CLT panels. Specimen details
CL5/145/2900
Mean Std COV CL5/145/4020
Mean Std COV
Test No
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
Experimental bending stiffness, EI 1011 (Nmm2) EIml
EImg with G = 1
EImg with G = 533
9.81 9.88 9.82 9.94 9.90 9.98 10.42 10.33 9.86 9.68 9.96 0.23 0.02 9.90 10.17 9.50 10.05 9.77 9.75 9.79 9.81 9.71 9.74 9.82 0.19 0.02
9.69 9.60 9.68 9.86 9.56 9.83 10.23 10.22 9.75 9.44 9.79 0.26 0.03 9.78 9.75 9.38 9.80 9.65 9.48 9.67 9.38 9.60 9.41 9.59 0.17 0.02
9.97 9.87 9.95 10.14 9.82 10.11 10.53 10.53 10.03 9.70 10.07 0.28 0.03 9.93 9.90 9.51 9.95 9.79 9.62 9.81 9.52 9.75 9.55 9.73 0.17 0.02
of the five-layer CLT panel was found to be 75% higher than that of the three-layer CLT panel with same span to thickness ratio. This indicates that the increasing thickness of the CLT panel will provide a higher shear load resistance capacity. The typical most common failure modes (i.e. rolling shear, delamination and finger joint failures) are shown in Fig. 7 and highlight that the shear is the dominant failure mode. The detailed test results inclusive of the experimental stiffness, maximum shear force, theoretical shear stress and failure modes are listed in Table 5. The highest stiffness
Mmax (kNm)
fb (MPa)
Failure mode
51.04 46.54 60.90 50.95 50.90 53.55 46.11 41.97 52.47 37.47 49.19 6.52 0.13 38.71 46.21 31.85 37.96 47.48 38.00 60.83 52.81 50.19 50.37 45.44 8.72 0.19
30.18 27.32 35.98 29.73 29.83 31.11 25.68 23.57 30.88 22.46 28.67 4.01 0.14 22.68 26.36 19.44 21.91 28.19 22.61 36.03 31.21 29.98 30.01 26.84 5.16 0.19
Tension Tension Tension Tension Tensile & Delamination Tension Tensile & Delamination Tension Tension Tension
Tension Tension Tensile & Delamination Tension Tension Tension Tension Tension Tensile & Delamination Tension
was observed in the five-layer CLT panels, namely 14.03E + 11 Nmm2 using global displacement with an infinite G, and 14.52E + 11 Nmm2 using local displacement. It was observed that increasing the thickness results in a higher stiffness (Table 5). The shear stress of each tested sample was derived from the layered beam theory, Gamma beam theory and shear analogy theory. Only the results for the specimens which exhibited shear failure were used to derive the maximum shear stress. The maximum shear stress for the five-layer CLT panels obtained from the
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
(a) CL3/105/2100
(b) CL3/105/2940
(c) CL5/145/2900
(d) CL5/145/4020
7
Fig. 5. Comparison between the experimental and theoretical bending stiffness of the tested CLT panels: (a) CL3/105/2100; (b) CL3/105/2940; (c) CL5/145/2900; (c) CL5/145/ 4020.
Fig. 6. Average force-displacement response for the shear test samples: (a) global force-displacement; (b) local force-displacement.
aforementioned theories were similar (Table 5). However, in the three-layer CLT panel, the maximum shear stress obtained from the layered beam theory and shear analogy theory were similar but slightly different to that obtained from the Gamma beam theory. The maximum shear stress obtained from the Gamma beam theory was about 5% less than that of from both layered beam and shear analogy theory in the three-layer CLT panel. The average
maximum shear stress of 1.89 N/mm2 and 1.82 N/mm2 were obtained for the three-layer and five-layer CLT panels, respectively. This indicates that an increasing thickness of the CLT panel results in a lower shear strength. The rolling shear stress ranged from 1.55 N/mm2 to 2.22 N/mm2, which are similar to those obtained by Li [28], who observed a variation in the mean rolling shear strength between 1.97 N/mm2 and 2.45 N/mm2. Li [28] also
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Fig. 7. The most common failure modes from the shear tests: (a) rolling shear failure; (b) delamination failure; (c) finger joint failure.
Table 5 Shear test results of three and five-layer CLT panels. Specimen details
CL3/105/ 1260
Mean Std COV CL5/145/ 1740
Mean Std COV
Test No
Experimental stiffness, EI 1011 (Nmm2) EIml
EImg with G =1
EImg with G = 533
T1 T2 T3 T4 T5 T6
3.88 3.87 4.09 4.00 3.91 3.93 3.97 4.04 3.98 3.96 0.08 0.02 14.52 14.29 14.12 14.42 14.09 13.96
3.70 3.74 3.79 3.81 3.78 3.85 3.75 3.74 3.71 3.76 0.05 0.01 14.03 14.00 13.97 13.94 13.91 13.78
3.94 3.99 4.04 4.06 4.02 4.11 3.99 3.98 3.95 4.01 0.05 0.01 15.11 15.09 15.06 15.03 14.98 14.84
T7 T8
14.23 14.34
14.03 13.83
T9 T10
14.45 14.38 14.28 0.23 0.02
T1 T2 T3 T4 T5 T6 T7 T8 T9
Max. shear force (kN)
Max. shear stress (N/mm2)
Failure mode
Layered Beam Theory
Gamma Beam Theory
Shear Analogy Theory
86 74 77 83 71 79 79 71 88 78.67 6.08 0.08 114 108 108 111 109 100
2.18 1.88 1.96 2.12 1.80 2.01 2.02 1.80 2.22 2.00 0.15 0.08 1.89 1.80 1.79 1.85 1.81 1.66
2.06 1.78 1.85 2.01 1.70 1.90 1.91 1.70 2.10 1.89 0.15 0.08 1.90 1.80 1.79 1.85 1.81 1.66
2.17 1.88 1.95 2.12 1.80 2.00 2.02 1.80 2.22 2.00 0.15 0.08 1.90 1.80 1.80 1.85 1.81 1.66
15.13 14.89
108 94
1.79 1.55
1.79 1.55
1.79 1.55
13.87 14.03
14.94 15.12
110 99
1.82 1.64
1.83 1.64
1.83 1.64
13.94 0.09 0.01
15.02 0.10 0.01
106.20 6.48 0.06
1.76 0.11 0.06
1.76 0.11 0.06
1.76 0.11 0.06
observed a similar trend, whereby thicker CLT panels exhibited lower shear strengths.
4. Finite element model 4.1. FEM development and validation A 3D FEM for the CLT panel under four point-bending was developed using ABAQUS [29] as shown in Fig. 8 and validated by experimental test data. This model aims to predict the bending and shear stiffness, and strength of the CLT panels. The incompatible mode eight-node brick element (C3D8I) was employed to model the lumber and this element can remove the shear locking and reduce the volumetric locking. Different layers of timber in the CLT panel were simulated by changing the material orientation of the lumber. A linear-elastic orthotropic material model implemented in ABAQUS [28] was used to simulate the mechanical behaviour of each lumber. The elastic and shear moduli of each layer (Table 1) were obtained from the manufacturer’s specification [21]. The Poisson’s
Shear Shear Shear Shear Shear Shear Shear Shear Shear
Shear Shear Shear Shear Shear Shear & delamination Shear Shear & delamination Shear Shear & delamination
ratios (i.e. mxz = 0.18, mxy = 0.2, myz = 0.21) were obtained from He, et al. [4]. Cohesive contact was used to simulate the mechanical behaviour of the adhesive between each layer and the cohesive properties were obtained from an experimental study by Oktavianus et al. [30]. The axial tensile capacity and the Young’s modulus of the adhesive is 1.8 MPa and 7.8 MPa, respectively. The fracture energy (Gf) was taken as 0.9 N/mm. Two numerical models were created for each type of CLT panel to predict the shear (CLT3/105/1260 and CLT5/145/1740) and bending (CLT3/105/2940 and CLT5/145/4020) behaviour. The average applied force versus the average global displacement as well as peak loads were obtained from the FEM analysis and compared with the experimental test results. Figs. 9 and 10 show the comparison of the average force versus global displacement relationship for the three and five-layer CLT panels with the experimental test results and the FEMs. Fig. 9 shows the load-displacement curve for the two bending samples CLT3/105/1260 and CLT3/105/2940. It can be observed that the ultimate load predicted by the FEM for the CLT3/105/1260 panel was 126.3 kN, approximately 2% less than that obtained from the
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
9
Fig. 8. The deformed and stress distribution of the CLT panel.
Fig. 9. Comparison of the force-displacement curves between the FEMs and experimental tests for CLT3/105/1260 and CLT3/105/2940.
experimental test (i.e. 128.5 kN). In the FEM of CLT3/105/2940, the ultimate load was 42.6 kN, which is about 15% lower than the ultimate load obtained from the experiment. Material non-linearity between the experiments and FEM might be the reason for the discrepancy in the peak load of CLT panel. Fig. 10 illustrates the load-displacement curves of CLT5/145/1740 and CLT5/145/4020 obtained from the FEM and average results from the experiments. The ultimate loads predicted by the FEMs in both specimens CLT5/145/1740 and CLT5/145/4020 were less than those obtained from the experimental results. However, the failure point of FEM was similar to the initial failure point of the experiments. The comparison between the FEM and experimental results of for samples CLT3/105/1260, CLT3/105/2940, CLT5/145/1740 and
CLT5/145/4020 show that the developed FEM models successfully captured the failure load and stiffness of the CLT panels, with maximum about 15% variations. Figs. 9 and 10 also show that the FEM captures the behavior of the CLT panel accurately in the linearelastic regime and predicts the initial failure point that was observed in the experiment. 4.2. Prediction of bending and shear stiffness The validated FEM models with three and five-layer CLT panels were used to predict the bending and shear strength of the sevenlayer CLT panels of the same span to thickness ratios of 12 and 28 as tested (i.e. CLT7/200/2400 and CLT7/200/5600). The CLT7 sample is made up of four 35 mm thick layers from timber grade
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Fig. 10. Comparison of the force-displacement curves between the FEMs and experimental tests for CLT5/145/1740 and CLT5/145/4020.
XLG1 and three 20 mm thick layers from timber grade XLG2. The FEM model predicted the ultimate loads of 116 kN and 216 kN for the CLT7/200/5600 and CLT7/200/2400, respectively (Fig. 11). The force–displacement relationship of the seven-layer CLT panels as shown in Fig. 11 was found to exhibits similar behaviour to that of the three and five-layer CLT panels (i.e. Figs. 9 and 10). This confirms that the numerical results of the seven-layers CLT panel are reliable.
The global and local stiffness of the CLT panels were derived from the FEM results and compared with that obtained from the experimental results (Table 6). Table 6 shows that the local and global stiffness of the CLT panels derived from the FEM was varied up to 20%, when compared with those obtained from the experimental tests. This variation could be attributed to the material non-linearity and flexibility in the supports and imperfect bonding between the layers of the CLT panel in the experimental test. In the
Fig. 11. The FEM predicted load-deformation curves for CLT7/200/2400 and CLT7/200/5600.
11
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215 Table 6 Comparison between the global and local stiffness of the CLT panels obtained from the FEMs and experimental tests. Specimen details
CL3/105/1260 CL5/145/1740 CL7/200/2400 CL3/105/2940 CL5/145/4020 CL7/200/5600
Average Experimental stiffness, EI 1011 (N. mm2)
FEM stiffness, EI 1011 (N.mm2)
EIml
EImg with G = 533
EIml
EImg with G = 533
3.96 14.28 – 5.14 9.82 –
4.01 15.02 – 4.75 9.73 –
4.14 13.30 27.90 4.70 11.80 29.11
3.47 12.90 21.40 4.52 11.01 27.31
FEM, the material properties were same for all the CLT panels, there is no flexibility at the supports and the bonding between each layer was assumed to be cohesive contact.
5. Conclusions The bending and shear behavior of Australian Radiata pine CLT have been studied using laboratory testing, and theoretical and FEM analysis. Based on these analyses, the following conclusions can be drawn: The average ultimate load (Fmax) for the five-layer CLT panel was about 10% higher than that of the three-layer CLT panel with the same span to thickness ratio. This indicates that increasing thickness of CLT panel had minimal influence on the load resistance capacity. The average bending strength of the five-layer CLT panel was approximately 15% higher than that of the three-layer CLT panel with a similar span to thickness ratio. This indicates that the influence of the panel thickness on the bending strength of the outermost layer is not significant. The average experimental EImg with an infinite shear modulus was about 18% higher than the theoretical bending stiffness in the three-layer CLT panel, whilst in the five-layer CLT panel, it was about 5% higher when compared to the theoretical bending stiffness. This result indicates that the theoretical prediction of EImg with an infinite shear modulus is possibly more accurate with an increased number of layers in the CLT panel. The average maximum shear stresses of 1.89 N/mm2 and 1.82 N/mm2 were obtained for the three and five-layer CLT panels, respectively. This indicates that shear strength decreases with increasing CLT panel thickness. The FEM analysis in this study can be an effective alternative to the experiments carried out to predict the bending and shear strength and stiffness of CLT panels. With proper input parameters, these methods could also be used to predict the failure modes of the CLT panel under different loading conditions. The experimental performance of Australian Radiata pine CLT panels manufactured by XLAM exceeded all theoretically calculated strengths and stiffnesses based on feedstock specifications and existing theoretical models. The results presented in this paper can provide a basis for the use of Australian made radiata pine CLT panels for structural applications.
Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
Acknowledgement This work was funded by the Australian Research Council (ARC) Centre for Advanced Manufacturing of Prefabricated Housing [Grant ID: IC150100023]. The authors also acknowledge Xlam Australia for supplying the CLT panels.
References [1] S. Navaratnam, T. Ngo, T. Gunawardena, D. Henderson, Performance review of prefabricated building systems and future research in Australia, Buildings 9 (2019) 38. [2] T. Gunawardena, T. Ngo, P. Mendis, J. Alfano, Innovative Flexible Structural System Using Prefabricated Modules, J. Archit. Eng. 22 (4) (2016), https://doi. org/10.1061/(ASCE)AE.1943-5568.0000214. In this issue. [3] T. Ngo, K.T.Q. Nguyen, P. Tran, Heat release and flame propagation in prefabricated modular unit with GFRP composite facades, J. Build. Simu. 9 (5) (2016) 607–616, https://doi.org/10.1007/s12273-016-0294-3. In this issue. [4] M.J. He, X.F. Sun, Z. Li, Bending and compressive properties of cross-laminated timber (CLT) panels made from Canadian hemlock, Constr. Build. Mater. 185 (2018) 175–183. [5] K.T.Q. Nguyen,, P. Weerasinghe, P. Mendis, T. Ngo, J. Barnett, Performance of modern building facades in fire: a comprehensive review, Electron. J. Struct. Eng. 16 (1) (2016) 69–86. In this issue. [6] A. Robertson, A comparative life cycle assessment of mid-rise office building construction alternatives: laminated timber or reinforced concrete Master’s Thesis, University of Toronto, 2007. [7] Architecture & Design, Forte by Lend Lease. Retrieved from Architecture & Design: https://www.architectureanddesign.com.au/projects/multiresidential/forte-by-lend-lease, (23.10.2014.). [8] I. Block, Cross-laminated-timber housing in east London offers the future of low carbon construction, Dezeen (2017). [9] D.P. Hindman, J.C. Bouldin, Mechanical properties of southern pine crosslaminated timber, J. Mater. Civ. Eng. 27 (9) (2015) 1–7. [10] H.-M. Park, M. Fushitani, K. Sato, T. Kubo, H.-S. Byeon, Bending creep performances of three-ply cross-laminated woods made with five species, J. Wood Sci. 52 (3) (2006) 220–229. [11] I.P. Christovasilis, M. Brunetti, M. Follesa, M. Nocetti, D. Vassallo, Evaluation of the mechanical properties of cross-laminated timber with elementary beam theories, Constr. Build. Mater. 122 (2016) 202–213. [12] A.L. Fortune, P. Quenneville, Feasibility study of New Zealand radiata pine cross-lam, in: S. Setunge (Ed.), Incorporating Sustainable Practice in Mechanics and Structures of Materials, CRC Press, 2010. [13] Y. Li, F. Lam, M. Li, R. Foschi, Duration-of-load effect on the rolling shear strength of cross laminated timber: 1. Duration-of-load tests and damage accumulation model, World conference on timber engineering (WCTE), Quebec City, Canada, August 10–14, 2014. [14] K.S. Sikora, D.O. McPolin, A.M. Harte, Effects of the thickness of crosslaminated timber (CLT) panels made from Irish Sitka spruce on mechanical performance in bending and shear, Constr. Build. Mater. 116 (2016) 141–150. [15] Olivier Perret, Arthur Lebée, Cyril Douthe, Karam Sab, Experimental determination of the equivalent-layer shear stiffness of CLT through fourpoint bending of sandwich beams, Construct. Build. Mater. Elsevier 186 (2018) 1132–1143. [16] Conan O’Ceallaigh, Karol Sikora, Annette Harte, The influence of panel lay-up on the characteristic bending and rolling shear strength of CLT, Buildings 8 (9) (2018) 114, https://doi.org/10.3390/buildings8090114. [17] ANSI/APA. PRG 320 Standard for Performance-Rated Cross-Laminated Timber. APA – The Engineered Wood Association. Tacoma, WA, 2012. [18] K. Lewis, R. Shrestha, K.I. Crews, Introduction to cross laminated timber and development of design procedures for Australia and New Zealand, in: S.T. Smith (Ed.). 23rd Australasian Conference on the Mechanics of Structures and Materials (ACMSM23), Byron Bay, Australia, 9–12 December 2014. [19] Woodco. 2012. Forest and Wood Products Industry Strategic Action Plan. Wood Council of New Zealand. http://woodco.org.nz/images/stories/pdfs/
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215 Strategic.pdf XLam NZ, 2013. Cross Laminated Timber Design Guide. XLam NZ, Nelson. A. Iqbal, Cross-laminated Timber for Building Structures. BRANZ Study Report SR336, BRANZ Ltd, Judgeford, New Zealand, 2015. XLam, Australian Cross Laminated Timber Panel Structural Guide, XLam Australia Design Guide V1, Australian Edition, July 2017. I.S. EN 16351:2015, Timber Structures – Cross Laminated Timber Requirements, CEN, 2015. S. Navaratnam, T. Ngo, S. Linforth, P.B. Christopher, The Use of Digital Image Correlation for Identifying Failure Characteristics of Cross-Laminated Timber Under Transverse Loading, 2019, The paper has been submitted to Measurement. EN 408: 2012-07, Timber structures – Structural Timber and Glued Laminated Timber-Determination of some Physical and Mechanical Properties, European Standard, European Committee for standardization.
[25] Standards Australia. Timber structures Part 1: Design methods, AS 1720.1, Standard Australia, Sydney, Australia, 2010. [26] E. Karacabeyli, B. Douglas, CLT Handbook. doi: 10.1139/cjfas-20120496, 2013. [27] EN 1995: 2008-06 Eurocode 5: Design of timber structures-part 1-1: GeneralCommon rules and rules for buildings, European Standard, European Committee for Standardization. [28] M. Li, Evaluating rolling shear strength properties of cross-laminated timber by short-span bending tests and modified planar shear tests, J. Wood Sci. 63 (4) (2017) 331–337. [29] ABAQUS version 6.12-3 [Computer software]. Dassault Systems Simulia, Providence, RI. [30] Y. Oktavianus, K.S. Kristombu Baduge, K. Orlowski, P. Mendis, Structural behaviour of prefabricated load bearing braced composite timber wall system, Eng. Struct. 176 (2018) 555–568.
Contents lists available at ScienceDirect
Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat
Bending and shear performance of Australian Radiata pine cross-laminated timber S. Navaratnam a,⇑, P.B. Christopher b, T. Ngo b,⇑, T.V. Le b a b
School of Engineering, RMIT University, Building 10| Level 13| Room 17A, 376-392 Swanston Street, Melbourne VIC 3001, Australia Department of Infrastructure Engineering, The University of Melbourne, Australia
h i g h l i g h t s Bending and shear behavior of Australian radiata pine cross-laminated timber. FEM analysis to predict bending and shear strength of CLT. Failure modes of CLT under out-of-plane bending and shear. Comparison of bending and shear stiffness of CLT between tests and theoretical.
a r t i c l e
i n f o
Article history: Received 15 April 2019 Received in revised form 7 October 2019 Accepted 10 October 2019
Keywords: Cross-laminated timber Bending strength Shear strength Failure mode Finite element model Radiata pine
a b s t r a c t Cross Laminated Timber (CLT) is increasingly being used in commercial and residential construction in Australia due to its inherent strength and sustainability credentials. Until recently, infrastructure building projects using CLT have been reliant on imported products from overseas manufacturers. There is now a viable Australian grown and fabricated CLT product from Radiata pine. This paper summarises the experimental results on the mechanical behaviour of Australian Radiata pine CLT panels in out-of-plane bending and shear. Three-layer 105 mm thick panels and five-layer 145 mm panels with three different spans were tested. The results demonstrate strong correlations to existing theoretical models and were used to validate the finite element model (FEM) developed in this research. The experimental results showed that the average bending stiffness of the CLT panels were marginally greater than theoretical values. The maximum bending strength for Radiata pine exceeded the characteristic strength of 14.0 MPa for grade XLG1 external laminas, with three-layer CLT samples averaging 28.7 MPa and five-layer CLT samples averaging 26.8 MPa. The maximum observed shear stresses ranged from 1.55 MPa to 2.18 MPa, which also exceeded the rolling shear characteristic strength of 1.2 MPa for the feedstock. The results also highlighted that the shear strength decreases with an increasing thickness of the CLT panel. Ó 2019 Elsevier Ltd. All rights reserved.
1. Introduction The construction industry demands continual improvements for building materials that are fast and safe to erect, cost effective and sustainable [1–3]. CLT shows potential to address many of these requirements in the industry, with a high degree of prefabrication, high in-plane and out-plane strength and stiffness, and good acoustic and thermal performance [4,5]. CLT is an engineered wood panel utilising at least three layers of timber boards glued
⇑ Corresponding authors. E-mail addresses: [email protected] (S. Navaratnam), [email protected]. au (P.B. Christopher), [email protected] (T. Ngo), [email protected]. au (T.V. Le). https://doi.org/10.1016/j.conbuildmat.2019.117215 0950-0618/Ó 2019 Elsevier Ltd. All rights reserved.
together in orthogonal directions. This combination of layers eliminates much of the natural variability and the influence of defects in timber products when compared to plain sawn timber. CLT panels are frequently fabricated using timber offcuts or lower grade feedstock from other structural timber productions, thereby having a lower embodied energy, greenhouse gas emissions and high carbon sequestration compared to typical concrete construction [6]. As a result of these advantages, CLT is being used more widely in building construction [4,7,8]. CLT panels are made from different timber species that depend on local resources such as Kiri, Katsura, Sugi, Hinoki, Buna spruce pine (Europe and Canada) and Radiata pine (Australia and New Zealand). Previous research has focused on the flexural and shear performance of a variety of timber feedstock for CLT panels
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
[4,9,10–16]. Hindman and Bouldin [9] performed four-point bending tests in order to evaluate the bending and shear properties of Southern pine CLT. Their results indicated an equal or higher shear resistance compared to the standard specified values of shear (i.e. GAeff = 23 MN/m) and bending (i.e. EIeff = 9.7 MNm) resistance [17]. Similarly, four-point bending tests were used to evaluate the stiffness and strength properties in the perpendicular and parallel directions to the principal plane of the CLT panels, which were made using spruce boards [11,14]. Sikora et al. [14] assessed the effect of the panel thickness of Irish Sitka spruce CLT and found that the bending and rolling shear strength decreased with an increasing panel thickness. The bending creep performance of engineered wood products has been assessed by Park, et al. [10], where CLT panels from five different timber species were tested and they found that the creep deformation perpendicular to the grain was decreased by cross laminating. Fortune and Quenneville [12], Lewis et al. [18], Woodco [19], Iqbal [20] all investigated the shear and bending performance of CLT panels fabricated from New Zealand Radiata pine. These previous studies found that the bending and shear strength, and stiffness of CLT panels were equal to or higher than the standard specifications. However, there is limited research that has assessed the bending and shear performance of CLT panels fabricated using Australian Radiata pine. This research investigates the bending and shear performance of Australian Radiata pine CLT panels using four-point bending tests and finite element (FE) analysis. Three different spans with two different thicknesses of CLT panels were tested to evaluate the bending and shear strength, and stiffness through experimental tests. The experimental results of this study also used to validate the FEM and predict the shear and bending performance of CLT panels with seven-layer.
2. Experimental tests and materials 2.1. Specimen and material details Experimental tests were performed to evaluate the strength and stiffness of CLT panels for both three and five-layer CLT build ups. The CLT panels were manufactured by XLam in 2018 with the three-layer 105 mm thick panel (CL3/105) made up with 35/35/35 mm layups, whilst the five-layer 145 mm thick panel (CL5/145) is made up of 35/20/35/20/35 mm layups. Each specimen was manufactured from 35 90 mm and 20 140 mm Australian Radiata pine timber and bonded with Purbond polyurethane adhesive without edge bonding. The thickness, width, modulus of elasticity (MOE), shear modulus and the lamination grades of the CLT test specimens are detailed in Table 1. Both the top and bottom outer layers are manufactured with XLG1 (comparable to F7) structural grade Radiata pine timber, while the inner layers are made with XLG2 (comparable to F5) structural grade Radiata pine timber [21]. The specified average MOE for the outer layers and inner layer is 8 GPa and 6 GPa, respectively. The average density of the CL3/105 panels is 492 kg/m3 with a coeffi-
cient of variation (COV) of 10%. The average density of the CL5/145 panels was 483 kg/m3 with a COV of 11%. 2.2. Test methods The CLT panels were tested using an MTS hydraulic actuator with a capacity of 500 kN and the load configurations were in accordance with I.S. EN 16351[22]. To derive the strength and stiffness properties, the standard I.S. EN 16,351 [22] was used, which specifies the span and number of test specimens based on the thickness (h) and width of the specimen. Four-point bending was applied over a span of 20 h and 28 h to derive the bending strength and stiffness properties, whereas 12 h spans were used to determine the shear (rolling) strength and stiffness perpendicular to the plane. Table 2 details dimensions of the specimens and the purpose of each test. The specimens were stored at 20 ± 2 °C and ~65% relative humidity for a minimum of 72 h prior to the experiments. The moisture content of each timber layer was measured and found to be 10.8% on average with a COV of 12%, which is above the manufacturer’s specification of not less than 8%. As specified by I.S. EN 16,351 [22] a constant displacement-controlled loading rate was applied and adjusted for each span and thickness of specimen to ensure the maximum load was achieved within 300 ± 120 s. Three laser transducers (LS1, LS2 and LS3) were used to measure and calculate the local and global displacements as shown in Fig. 1. The laser transducers LS1 and LS3 were positioned such that the central gauge length was five times that of the depth of the CLT panel. The global displacement was measured by the laser transducer LS2 at the mid-span of the test specimen. The local displacement was determined by deducting the average displacement obtained from LS1 and LS3 from the global displacement (LS2). Multiple cameras were used to assess the lateral and vertical movement of the panels, and capture the crack propagation through the thickness. The cameras were positioned near the supports, shear zones and the shear-free zone (Fig. 1). Digital image correlation (DIC) analysis was performed to assess the out-ofplane movement panel, in-plane strain and crack growth of CLT [23]. Based on this analysis, its noted that the lateral and vertical movement at the supports was negligible until 90% of the ultimate load was reached. Specifically, a sudden lateral displacement between 10 and 15 mm occurred at the supports during the shear test when the ultimate load was reached. This movement does not affect the global and local stiffness of the panel as the flexural rigidities were calculated between 10% and 40% of the ultimate load using Eqs. (1) and (2). 3. Results and analysis 3.1. Bending test results The global load-displacement curve for the ten CL3/105 samples tested over a span of 2940 mm is shown in Fig. 2. It can be
Table 1 CLT test specimen details and material properties. Specimen details Type
Lamination grade
Lamination thickness (mm)
Width (mm)
Average density (kg/m3) /COV (%)
CL3/105 CL5/145 Mechanical properties Lamination grade MOE (MPa)
XLG1/ XLG2/ XLG1 XLG1/ XLG2/ XLG1/ XLG2/ XLG1
35/35/35 35/20/35/20/35
520 520
492/10 483/11
Parallel to the grain Perpendicular to the grain Longitudinal (G0) Rolling shear (GRS)
XLG2 6000 200 400 40
XLG1 8000 267 533 53
Shear modulus (MPa)
3
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215 Table 2 CLT panel test specimens loaded perpendicular to the plane. Specimen label
Number of samples
Number of layers
Thickness (mm)
Span (mm)
Loading direction
Determined properties
CL3/105/1260 CL5/145/1740 CL3/105/2100 CL3/105/2940 CL5/145/2900 CL5/145/4020
9 10 10 10 10 10
3 5 3 3 5 5
105 145 105 105 145 145
1260 1740 2100 2940 2900 4020
Perpendicular Perpendicular Perpendicular Perpendicular Perpendicular Perpendicular
Shear strength & stiffness Shear strength & stiffness Bending strength & stiffness Bending strength & stiffness Bending strength & stiffness Bending strength & stiffness
to to to to to to
plane plane plane plane plane plane
Fig. 1. Testing set-up: (a) Experimental test set-up; (b) Schematic diagram and nomenclature.
observed that the ultimate load (Fmax) of each test specimen varies between 46.7 kN and 56.7 kN with a COV of 7%. Fig. 2 also shows that all samples exhibited brittle failure at the ultimate load. The average global and local force–displacement curves for each bending test sample (i.e. CL3/105/2100, CL3/105/2940, CL5/145/2900, CL5/145/4020) were obtained from the tests and are shown in Fig. 3. It can be observed that the average ultimate load (Fmax) for the five-layer CLT panel was about 10% higher than the threelayer CLT panel with the same span to thickness ratio. This indicates that increasing the thickness of the panel whilst maintaining the same span-to-thickness ratio does not improve its flexural capacity. Similar phenomenon was in agreement with the finding made by Sikora, et al. [14] for the CLT panels made from Irish Sitka spruce. The most common failure modes observed in the bending tests were tensile failure of the lower lamina, delamination between the lower laminas and tensile failure at the finger joints, as shown in Fig. 4. The global displacement which accounts for both shear and bending deformations was used together with the force for each sample to derive the global stiffness (EImg). In a similar way the local force–displacement relationship was used to calculate the local flexural stiffness (i.e. bending stiffness in the absence of shear deformation) in the absence of shear stresses (EIml). Based on EN
408 [24], the following equations were used to derive the global and local stiffness, which were similar to those used by Sikora et al. [14] and He et al. [4]: 2
EIml ¼
al1 ðF 2 F 1 Þ 16ðw2 w1 Þ
EImg ¼
3aL2 4a2 3a 1 48ðwF 22 w 5Gbh Þ F 1
ð1Þ
ð2Þ
where: a is the distance between the loading head and the nearest support; l1 (5 h) is equal to the gauge length for the local displacement measurement; L is the span (Fig. 1); F1 and F2 are the loads corresponding to 10% and 40% of the ultimate load (Fmax); w1 and w2 are the displacements corresponding to loads F1 and F2, respectively; G is the shear modulus; and b is the width of the panel (520 mm). Two global stiffnesses were determined by taking the shear modulus as both infinite and 533 N/mm2 (see Tables 3 and 4). The highest stiffness obtained for the three-layer CLT panels were calculated as 5.39E + 11 Nmm2 using the local displacement and 4.67E + 11 Nmm2 using the global displacement, with an infinite shear modulus. Both the local and global stiffnesses were higher in the fivelayer CLT panels, namely about 93% and 119%, respectively. These
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Force (kN)
4
CL3/105/2940_Average CL3/105/2940_T1 CL3/105/2940_T2 CL3/105/2940_T3 CL3/105/2940_T4 CL3/105/2940_T5 CL3/105/2940_T6 CL3/105/2940_T7 CL3/105/2940_T8 CL3/105/2940_T9 CL3/105/2940_T10
60 50 40 30 20
10 0 0
10
20
30
40 50 60 Global displacement (mm)
70
Fig. 2. Global force-displacement response for sample CL3/105/2940.
(a)Global displacement
(b)Local displacement
Fig. 3. Average force-displacement response for bending test samples: (a) global force-displacement; (b) local force-displacement.
results indicate that an increase in the thickness of the CLT panel by 40% will result in doubling of the bending stiffness. The bending strength (fb) of the CLT panels was obtained using the Shear Analogy theory which takes into consideration the shear deformation of the cross laminated layers and is tabulated in Tables 3 and 4. The flexural strength of the outer laminates (fb) of the CLT panel can be derived using Eq. (3).
fb ¼
M max E1 h=2 EIml
ð3Þ
where: Mmax is the maximum bending moment; E1 is the MOE of the outer layer; and EIml is the local stiffness obtained from Eq. (1). The highest fb recorded for the three-layer and five-layer panels were 33.45 MPa and 36.03 MPa, respectively. The average fb for the five-layer CLT panel (i.e. 26.84 MPa) was approximately 15% higher than that of the three-layer panel (i.e. 23.41 MPa) with the same span-to-thickness ratio. This indicates that increasing
the panel thickness does not improve the bending strength of the outermost layer for panels with the same span-to-thickness ratio. Further, the average fb values for the Australian radiata pine CLT are approximately between the bending properties of machine graded pine (MGP) 10 and 12 as specified in Australian Standard AS 1720.1 [25]. 3.2. Experimental and theoretical comparison of the bending stiffness Three methods were used to calculate the theoretical bending stiffness (EIeff) for each CLT panel; the Shear Analogy Theory, Modified Gamma Theory and Composite Theory (k method). Eqs. (4) and (5) as specified by Karacabeyli and Douglas [26] were used to calculate the effective bending stiffness based on the Shear Analogy theory (EIeff, shear) and Composite theory (EIeff,comp), respectively. Eq. (6) was used to determine the effective bending stiffness using the Modified Gamma theory, which does not con-
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
5
Fig. 4. The most common failure modes from the bending tests: a) tension failure; b) delamination; c) tension failure at the finger joints.
sider shear deformation and is founded on the Mechanically Jointed Beams theory EN 1995 [27].
EIeff ;shear ¼
n X i¼1
EIeff ;Comp ¼ Eeff :
EIeff ;Gamma ¼
n X hi Ei :Ai :zi 2 þ 12 i¼1 3
Ei :bi :
b:htot 12
n X
ð4Þ
3
ð5Þ 3
Ei :bi :
i¼1
hi 12
! þ
Xn i
ci :Ei :bi :hi :zi 2
ð6Þ
where: bi is the width of each layer; hi is the thickness of each layer; zi is the distance between the center point of each layer and the neutral axis; Ei is the MOE; htot is the thickness of the panel; and Eeff = 30 k1E, k1 is determined by the load configuration. The connection efficiency factor ci , which can be derived from Equation (7) below, is non-zero only for longitudinal layers and equal to unity for the middle layer.
ci ¼ 1 þ
p2 :Ei :bi :hi
L2 eff : G:bj =hj
!1 ð7Þ
where: Leff is the effective length of the beam; and j refers to the transverse layer connecting the ith layer with the central layer.
Fig. 5 shows a comparison between the theoretical and experimental bending stiffness. The theoretical values of EIeff,shear, EIeff,comp and EIeff,Gamma were found to yield similar estimations of the bending stiffness. The experimentally determined global bending stiffness EImg was consistently less than the experimentally determined local bending stiffness EIml, however both experimental values were found to be higher than any of the theoretical bending stiffnesses. Similar results were reported by Sikora et al. [14] and He et al. [4]. Fig. 5 also shows a comparison between the theoretical and experimental bending stiffness, and illustrates that the EImg with an infinite shear modulus gives less differences than EIml and EImg with a finite shear modulus of 533 MPa. This indicates that the theoretical equations can be accurately predict the global bending stiffness with shear and bending deformation. The average EImg with an infinite shear modulus (4.53E + 11 Nmm2) was about 18% higher than the theoretical bending stiffness in the three-layer CLT panel, whilst in the five-layer CLT panel, it was about 5% higher when compared to the theoretical bending stiffness. This indicates that the theoretical prediction of EImg with an infinite shear modulus is possibly more accurate with an increased number of layers in the CLT panel. 3.3. Shear test results The average global and local force-displacement relationship from the shear tests are shown in Fig. 6. The ultimate load (Fmax)
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Table 3 Bending test results for the tree-layer CLT panels. Specimen details
Test No
Experimental bending stiffness, EI 1011 (Nmm2) EIml
EImg with G =1
EImg with G = 533
CL3/105/2100
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
5.31 4.93 4.94 5.32 4.96 5.10 4.97 5.04 5.39 5.13 5.11 0.17 0.03 5.13 5.03 5.41 5.22 5.06 5.19 5.16 5.00 5.02 5.17 5.14 0.12 0.02
4.50 4.26 4.33 4.56 4.29 4.44 4.33 4.35 4.67 4.42 4.42 0.13 0.03 4.66 4.55 4.78 4.74 4.63 4.77 4.70 4.55 4.60 4.65 4.66 0.08 0.02
4.66 4.40 4.47 4.72 4.43 4.59 4.47 4.50 4.84 4.57 4.57 0.14 0.03 4.75 4.63 4.87 4.84 4.72 4.86 4.79 4.63 4.68 4.74 4.75 0.09 0.02
Mean Std COV CL3/105/2940
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
Mean Std COV
Mmax (kNm)
fb,0 (MPa)
Failure mode
38.21 34.73 24.16 27.58 31.35 40.55 34.52 31.23 31.00 30.13 32.35 4.84 0.15 27.18 31.59 31.58 26.06 26.54 27.66 29.50 29.33 26.58 30.45 28.65 2.11 0.07
30.20 29.56 20.53 21.78 26.52 33.45 29.19 26.02 24.17 24.69 26.61 4.02 0.15 22.27 26.37 24.49 20.97 22.02 22.39 23.99 24.66 22.26 24.73 23.41 1.67 0.07
Tensile & Delamination Tension Tension Tension Tension Tensile & Delamination Tensile & Delamination Tension Tension Tension
Tension Tension Tension Tensile & Delamination Tension Tension Tensile & Delamination Tension Tension Tension
Table 4 Bending test results for the five-layer CLT panels. Specimen details
CL5/145/2900
Mean Std COV CL5/145/4020
Mean Std COV
Test No
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
T1 T2 T3 T4 T5 T6 T7 T8 T9 T10
Experimental bending stiffness, EI 1011 (Nmm2) EIml
EImg with G = 1
EImg with G = 533
9.81 9.88 9.82 9.94 9.90 9.98 10.42 10.33 9.86 9.68 9.96 0.23 0.02 9.90 10.17 9.50 10.05 9.77 9.75 9.79 9.81 9.71 9.74 9.82 0.19 0.02
9.69 9.60 9.68 9.86 9.56 9.83 10.23 10.22 9.75 9.44 9.79 0.26 0.03 9.78 9.75 9.38 9.80 9.65 9.48 9.67 9.38 9.60 9.41 9.59 0.17 0.02
9.97 9.87 9.95 10.14 9.82 10.11 10.53 10.53 10.03 9.70 10.07 0.28 0.03 9.93 9.90 9.51 9.95 9.79 9.62 9.81 9.52 9.75 9.55 9.73 0.17 0.02
of the five-layer CLT panel was found to be 75% higher than that of the three-layer CLT panel with same span to thickness ratio. This indicates that the increasing thickness of the CLT panel will provide a higher shear load resistance capacity. The typical most common failure modes (i.e. rolling shear, delamination and finger joint failures) are shown in Fig. 7 and highlight that the shear is the dominant failure mode. The detailed test results inclusive of the experimental stiffness, maximum shear force, theoretical shear stress and failure modes are listed in Table 5. The highest stiffness
Mmax (kNm)
fb (MPa)
Failure mode
51.04 46.54 60.90 50.95 50.90 53.55 46.11 41.97 52.47 37.47 49.19 6.52 0.13 38.71 46.21 31.85 37.96 47.48 38.00 60.83 52.81 50.19 50.37 45.44 8.72 0.19
30.18 27.32 35.98 29.73 29.83 31.11 25.68 23.57 30.88 22.46 28.67 4.01 0.14 22.68 26.36 19.44 21.91 28.19 22.61 36.03 31.21 29.98 30.01 26.84 5.16 0.19
Tension Tension Tension Tension Tensile & Delamination Tension Tensile & Delamination Tension Tension Tension
Tension Tension Tensile & Delamination Tension Tension Tension Tension Tension Tensile & Delamination Tension
was observed in the five-layer CLT panels, namely 14.03E + 11 Nmm2 using global displacement with an infinite G, and 14.52E + 11 Nmm2 using local displacement. It was observed that increasing the thickness results in a higher stiffness (Table 5). The shear stress of each tested sample was derived from the layered beam theory, Gamma beam theory and shear analogy theory. Only the results for the specimens which exhibited shear failure were used to derive the maximum shear stress. The maximum shear stress for the five-layer CLT panels obtained from the
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
(a) CL3/105/2100
(b) CL3/105/2940
(c) CL5/145/2900
(d) CL5/145/4020
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Fig. 5. Comparison between the experimental and theoretical bending stiffness of the tested CLT panels: (a) CL3/105/2100; (b) CL3/105/2940; (c) CL5/145/2900; (c) CL5/145/ 4020.
Fig. 6. Average force-displacement response for the shear test samples: (a) global force-displacement; (b) local force-displacement.
aforementioned theories were similar (Table 5). However, in the three-layer CLT panel, the maximum shear stress obtained from the layered beam theory and shear analogy theory were similar but slightly different to that obtained from the Gamma beam theory. The maximum shear stress obtained from the Gamma beam theory was about 5% less than that of from both layered beam and shear analogy theory in the three-layer CLT panel. The average
maximum shear stress of 1.89 N/mm2 and 1.82 N/mm2 were obtained for the three-layer and five-layer CLT panels, respectively. This indicates that an increasing thickness of the CLT panel results in a lower shear strength. The rolling shear stress ranged from 1.55 N/mm2 to 2.22 N/mm2, which are similar to those obtained by Li [28], who observed a variation in the mean rolling shear strength between 1.97 N/mm2 and 2.45 N/mm2. Li [28] also
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Fig. 7. The most common failure modes from the shear tests: (a) rolling shear failure; (b) delamination failure; (c) finger joint failure.
Table 5 Shear test results of three and five-layer CLT panels. Specimen details
CL3/105/ 1260
Mean Std COV CL5/145/ 1740
Mean Std COV
Test No
Experimental stiffness, EI 1011 (Nmm2) EIml
EImg with G =1
EImg with G = 533
T1 T2 T3 T4 T5 T6
3.88 3.87 4.09 4.00 3.91 3.93 3.97 4.04 3.98 3.96 0.08 0.02 14.52 14.29 14.12 14.42 14.09 13.96
3.70 3.74 3.79 3.81 3.78 3.85 3.75 3.74 3.71 3.76 0.05 0.01 14.03 14.00 13.97 13.94 13.91 13.78
3.94 3.99 4.04 4.06 4.02 4.11 3.99 3.98 3.95 4.01 0.05 0.01 15.11 15.09 15.06 15.03 14.98 14.84
T7 T8
14.23 14.34
14.03 13.83
T9 T10
14.45 14.38 14.28 0.23 0.02
T1 T2 T3 T4 T5 T6 T7 T8 T9
Max. shear force (kN)
Max. shear stress (N/mm2)
Failure mode
Layered Beam Theory
Gamma Beam Theory
Shear Analogy Theory
86 74 77 83 71 79 79 71 88 78.67 6.08 0.08 114 108 108 111 109 100
2.18 1.88 1.96 2.12 1.80 2.01 2.02 1.80 2.22 2.00 0.15 0.08 1.89 1.80 1.79 1.85 1.81 1.66
2.06 1.78 1.85 2.01 1.70 1.90 1.91 1.70 2.10 1.89 0.15 0.08 1.90 1.80 1.79 1.85 1.81 1.66
2.17 1.88 1.95 2.12 1.80 2.00 2.02 1.80 2.22 2.00 0.15 0.08 1.90 1.80 1.80 1.85 1.81 1.66
15.13 14.89
108 94
1.79 1.55
1.79 1.55
1.79 1.55
13.87 14.03
14.94 15.12
110 99
1.82 1.64
1.83 1.64
1.83 1.64
13.94 0.09 0.01
15.02 0.10 0.01
106.20 6.48 0.06
1.76 0.11 0.06
1.76 0.11 0.06
1.76 0.11 0.06
observed a similar trend, whereby thicker CLT panels exhibited lower shear strengths.
4. Finite element model 4.1. FEM development and validation A 3D FEM for the CLT panel under four point-bending was developed using ABAQUS [29] as shown in Fig. 8 and validated by experimental test data. This model aims to predict the bending and shear stiffness, and strength of the CLT panels. The incompatible mode eight-node brick element (C3D8I) was employed to model the lumber and this element can remove the shear locking and reduce the volumetric locking. Different layers of timber in the CLT panel were simulated by changing the material orientation of the lumber. A linear-elastic orthotropic material model implemented in ABAQUS [28] was used to simulate the mechanical behaviour of each lumber. The elastic and shear moduli of each layer (Table 1) were obtained from the manufacturer’s specification [21]. The Poisson’s
Shear Shear Shear Shear Shear Shear Shear Shear Shear
Shear Shear Shear Shear Shear Shear & delamination Shear Shear & delamination Shear Shear & delamination
ratios (i.e. mxz = 0.18, mxy = 0.2, myz = 0.21) were obtained from He, et al. [4]. Cohesive contact was used to simulate the mechanical behaviour of the adhesive between each layer and the cohesive properties were obtained from an experimental study by Oktavianus et al. [30]. The axial tensile capacity and the Young’s modulus of the adhesive is 1.8 MPa and 7.8 MPa, respectively. The fracture energy (Gf) was taken as 0.9 N/mm. Two numerical models were created for each type of CLT panel to predict the shear (CLT3/105/1260 and CLT5/145/1740) and bending (CLT3/105/2940 and CLT5/145/4020) behaviour. The average applied force versus the average global displacement as well as peak loads were obtained from the FEM analysis and compared with the experimental test results. Figs. 9 and 10 show the comparison of the average force versus global displacement relationship for the three and five-layer CLT panels with the experimental test results and the FEMs. Fig. 9 shows the load-displacement curve for the two bending samples CLT3/105/1260 and CLT3/105/2940. It can be observed that the ultimate load predicted by the FEM for the CLT3/105/1260 panel was 126.3 kN, approximately 2% less than that obtained from the
S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
9
Fig. 8. The deformed and stress distribution of the CLT panel.
Fig. 9. Comparison of the force-displacement curves between the FEMs and experimental tests for CLT3/105/1260 and CLT3/105/2940.
experimental test (i.e. 128.5 kN). In the FEM of CLT3/105/2940, the ultimate load was 42.6 kN, which is about 15% lower than the ultimate load obtained from the experiment. Material non-linearity between the experiments and FEM might be the reason for the discrepancy in the peak load of CLT panel. Fig. 10 illustrates the load-displacement curves of CLT5/145/1740 and CLT5/145/4020 obtained from the FEM and average results from the experiments. The ultimate loads predicted by the FEMs in both specimens CLT5/145/1740 and CLT5/145/4020 were less than those obtained from the experimental results. However, the failure point of FEM was similar to the initial failure point of the experiments. The comparison between the FEM and experimental results of for samples CLT3/105/1260, CLT3/105/2940, CLT5/145/1740 and
CLT5/145/4020 show that the developed FEM models successfully captured the failure load and stiffness of the CLT panels, with maximum about 15% variations. Figs. 9 and 10 also show that the FEM captures the behavior of the CLT panel accurately in the linearelastic regime and predicts the initial failure point that was observed in the experiment. 4.2. Prediction of bending and shear stiffness The validated FEM models with three and five-layer CLT panels were used to predict the bending and shear strength of the sevenlayer CLT panels of the same span to thickness ratios of 12 and 28 as tested (i.e. CLT7/200/2400 and CLT7/200/5600). The CLT7 sample is made up of four 35 mm thick layers from timber grade
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215
Fig. 10. Comparison of the force-displacement curves between the FEMs and experimental tests for CLT5/145/1740 and CLT5/145/4020.
XLG1 and three 20 mm thick layers from timber grade XLG2. The FEM model predicted the ultimate loads of 116 kN and 216 kN for the CLT7/200/5600 and CLT7/200/2400, respectively (Fig. 11). The force–displacement relationship of the seven-layer CLT panels as shown in Fig. 11 was found to exhibits similar behaviour to that of the three and five-layer CLT panels (i.e. Figs. 9 and 10). This confirms that the numerical results of the seven-layers CLT panel are reliable.
The global and local stiffness of the CLT panels were derived from the FEM results and compared with that obtained from the experimental results (Table 6). Table 6 shows that the local and global stiffness of the CLT panels derived from the FEM was varied up to 20%, when compared with those obtained from the experimental tests. This variation could be attributed to the material non-linearity and flexibility in the supports and imperfect bonding between the layers of the CLT panel in the experimental test. In the
Fig. 11. The FEM predicted load-deformation curves for CLT7/200/2400 and CLT7/200/5600.
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S. Navaratnam et al. / Construction and Building Materials 232 (2020) 117215 Table 6 Comparison between the global and local stiffness of the CLT panels obtained from the FEMs and experimental tests. Specimen details
CL3/105/1260 CL5/145/1740 CL7/200/2400 CL3/105/2940 CL5/145/4020 CL7/200/5600
Average Experimental stiffness, EI 1011 (N. mm2)
FEM stiffness, EI 1011 (N.mm2)
EIml
EImg with G = 533
EIml
EImg with G = 533
3.96 14.28 – 5.14 9.82 –
4.01 15.02 – 4.75 9.73 –
4.14 13.30 27.90 4.70 11.80 29.11
3.47 12.90 21.40 4.52 11.01 27.31
FEM, the material properties were same for all the CLT panels, there is no flexibility at the supports and the bonding between each layer was assumed to be cohesive contact.
5. Conclusions The bending and shear behavior of Australian Radiata pine CLT have been studied using laboratory testing, and theoretical and FEM analysis. Based on these analyses, the following conclusions can be drawn: The average ultimate load (Fmax) for the five-layer CLT panel was about 10% higher than that of the three-layer CLT panel with the same span to thickness ratio. This indicates that increasing thickness of CLT panel had minimal influence on the load resistance capacity. The average bending strength of the five-layer CLT panel was approximately 15% higher than that of the three-layer CLT panel with a similar span to thickness ratio. This indicates that the influence of the panel thickness on the bending strength of the outermost layer is not significant. The average experimental EImg with an infinite shear modulus was about 18% higher than the theoretical bending stiffness in the three-layer CLT panel, whilst in the five-layer CLT panel, it was about 5% higher when compared to the theoretical bending stiffness. This result indicates that the theoretical prediction of EImg with an infinite shear modulus is possibly more accurate with an increased number of layers in the CLT panel. The average maximum shear stresses of 1.89 N/mm2 and 1.82 N/mm2 were obtained for the three and five-layer CLT panels, respectively. This indicates that shear strength decreases with increasing CLT panel thickness. The FEM analysis in this study can be an effective alternative to the experiments carried out to predict the bending and shear strength and stiffness of CLT panels. With proper input parameters, these methods could also be used to predict the failure modes of the CLT panel under different loading conditions. The experimental performance of Australian Radiata pine CLT panels manufactured by XLAM exceeded all theoretically calculated strengths and stiffnesses based on feedstock specifications and existing theoretical models. The results presented in this paper can provide a basis for the use of Australian made radiata pine CLT panels for structural applications.
Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
Acknowledgement This work was funded by the Australian Research Council (ARC) Centre for Advanced Manufacturing of Prefabricated Housing [Grant ID: IC150100023]. The authors also acknowledge Xlam Australia for supplying the CLT panels.
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