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Structural performance of hybrid SPFs-LSL cross-laminated timber panels
Construction and Building Materials 149 (2017) 156–163
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Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat
Structural performance of hybrid SPFs-LSL cross-laminated timber panels William G. Davids a,⇑, Nicholas Willey a, Roberto Lopez-Anido a, Stephen Shaler b, Douglas Gardner b, Russell Edgar c, Mehdi Tajvidi b a b c
Department of Civil and Environmental Engineering, University of Maine, 5711 Boardman Hall, Orono, MA 04469-5711, USA School of Forest Resources, University of Maine, 5755 Nutting Hall, Orono, MA 04469-5755, USA Advanced Structures and Composites Center, University of Maine, 35 Flagstaff Road, Orono, MA 04469-5793, USA
h i g h l i g h t s Hybrid CLT with laminated strand lumber (LSL) tested under out of plane loads. LSL core eliminated shear failure in 3-layer panels. CLT panel bending strength increased by 23% through inclusion of LSL core.
a r t i c l e
i n f o
Article history: Received 6 December 2016 Received in revised form 10 May 2017 Accepted 13 May 2017
Keywords: Cross-laminated timber Laminated strand lumber Wood shear strength Wood flexural strength Structural testing
a b s t r a c t The bending and shear performance of hybrid cross-laminated timber (CLT) panels made from SprucePine-Fir (South) (SPFs) and laminated strand lumber (LSL) are examined. Four configurations of threelayer CLT were fabricated: all-SPFs control specimens, all-LSL specimens, hybrid specimens with SPFs faces and an LSL core, and hybrid specimens with LSL faces and an SPFs core. Bending tests were conducted to assess flexural strength and stiffness. Additionally, three-point bending tests were performed to assess shear performance. The incorporation of LSL in the core of CLT panels increased mean panel bending stress at failure by 23% through mitigation of rolling shear failure. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Originally introduced in Austria and Germany in the mid-1990s, cross-laminated timber (CLT) has become an increasingly popular alternative for multi-story timber construction in Europe [1]. CLT has recently garnered interest in North America with the establishment of several CLT and nail-laminated timber plants in Canada and the United States. CLT panels are suitable for use in walls, floors and roofs, and are typically fabricated from an odd number of flat-wise layers of solid-sawn lumber placed in alternating 90 degree directions. In the majority of cases, individual layers of boards are adhesively bonded although nail- and screwlaminated CLT is also produced. Alternative forms of CLT have been considered including placing laminations at ±45 degrees as well as hollow, box-based systems [2]. Compared to typical concrete con⇑ Corresponding author. E-mail address: [email protected] (W.G. Davids). http://dx.doi.org/10.1016/j.conbuildmat.2017.05.131 0950-0618/Ó 2017 Elsevier Ltd. All rights reserved.
struction, CLT structures are lightweight, sequester more carbon, possess better thermal insulation properties, and are more rapidly erected [3]. Research on the structural performance of CLT can be separated into the broad categories of seismic behavior [4–7], fire resistance [8,9], and determination of the mechanical properties of CLT. A number of studies have focused on the determination of CLT mechanical properties in flexure and shear, which are primary design properties for panels subjected to out-of-plane loading [10–15]. Sikora et al. [10] present a current and thorough review of the existing literature on this topic. Others have focused on CLT mechanical response due to in-plane loading [16–18]. As with plywood, an issue which can limit the capacity of CLT subjected to out-of-plane loading is failure in perpendicular-tograin shear, commonly called rolling shear. Rolling shear also contributes to deflections of CLT panels. Because of its significance, several investigations have considered rolling shear properties and failure mechanisms. Zhou et al. [19] examined the effect of
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rolling shear deformations in 3-layer, black spruce CLT, measuring rolling shear modulus and conducting three-point bending tests of CLT specimens. Zhou et al. [19] also proposed a deflection adjustment factor to account for rolling shear deformations, and also concluded that bending specimen width did not significantly affect apparent elastic modulus and apparent shear modulus. Li et al. [20] implemented a torsional test for evaluating rolling shear strength in CLT, observing that thinner cross-layers tended to have higher rolling shear strengths. Hochreiner et al. [14] studied CLT plates subjected to concentrated loads and examined the evolution of rolling shear failure modes by tracking fracture development and load-deformation history using digital image correlation. Li and Lam [21] experimentally assessed rolling shear damage accumulation in CLT attributable to load cycling, and calibrated a damage accumulation model that can be used for future studies on duration-of-load behavior of CLT under rolling shear. Noting the significance of rolling shear on CLT structural performance, Aicher et al. [22] assessed the rolling shear modulus and strength of European beech, which typically has much better rolling shear properties than softwoods normally used in CLT construction, concluding that the use of beech in CLT cross-layers could be beneficial for CLT strength. Wang et al. [23] assessed the use of laminated strand lumber (LSL) in both cross-layers and face layers of hybrid CLT panels, demonstrating increased flexural capacities relative to conventional all-softwood CLT. The literature indicates that rolling shear failure can be a limiting factor for the strength of CLT subjected to out-of-plane loading. The focus of the research reported in this paper was the structural assessment of hybrid CLT panels made from LSL and softwood lumber with the objective of increasing strength by mitigating rolling shear failures in the core layer. LSL is an engineered composite lumber that is made from approximately 300 mm long strands of fast-growing species (often aspen or poplar) that are bonded and densified during manufacture and oriented with the long axis of the structural member. LSL typically possesses good dimensional stability and very predictable strength and stiffness values compared to solid-sawn lumber. Additionally, the authors are aware of no published experimental research specifically examining the use of Northeastern U.S. Spruce-Pine-Fir (South) (SPFs) lumber in CLT. SPFs is an economically significant group of lumber species harvested in the United States that includes Eastern Spruces, Balsam Fir, Red Pine, Jack Pine, Englemann Spruce, Lodgepole Pine, Sitka Spruce and Norway Spruce. All species are subject to the same grading rules and have the same design values. CLT panel production using SPFs harvested and milled in the Northeastern US may become an important new market for lumber producers in the United States as CLT markets grow. The research reported in this paper includes characterization of the SPFs and LSL lumber used for CLT manufacturing, assessment of the bond between SPFs and LSL using a polyurethane adhesive, and testing to assess both major-axis flexural and shear strength and stiffness.
2. Materials and methods Materials were 38 mm 184 mm 3 m kiln-dried SPFs No. 2, 38 mm 184 mm 3 m grade 1.35E LSL boards (without wax coating on the board edges), and Henkel PURBOND HB E452 single-component polyurethane adhesive. The No. 2 grade of SPFs is a standard grading category corresponding to specific stiffness and strength design values in the US codes, and is a commonly produced grade of SPFs lumber. The SPFs lumber was procured in bulk quantities from Pleasant River Lumber in Dover-Foxcroft, Maine, USA, and as discussed later, a small percentage of the SPFs was No. 1, a higher grade with higher design values. The LSL was provided by Louisiana-Pacific Corporation’s plant in Houlton, Maine, USA. The designation ‘‘1.35E LSL” refers to a specific grade of LSL that has a nominal elastic modulus of 9310 MPa. The 1.35E LSL was selected as opposed to a higher grade – 1.55E and 1.75E grades with moduli of 10,700 MPA and 12,070 MPA are also available – because its flexural strength and stiffness were expected to be similar to the SPFs.
Two, three-layer CLT panels were laid up for each of four configurations as described in Table 1. The CLT panels were made from continuous boards as opposed to the finger-jointed lumber typically used in commercially fabricated CLT panels.
2.1. Lumber characterization and preparation Each SPFs and LSL board was first E-rated using a Metriguard 340 E-Computer, its moisture content (MC) taken with a Delhorst J2000 pin moisture meter, and its dimensions measured and density calculated. In the E-rating process, the board is placed flatwise on two supports, one of which contains a small load cell. The board is struck at mid-span with a hammer, and the dynamic load cell readings are used to compute a dynamic modulus of elasticity (MOE). A total of over 900 SPFs and over 700 LSL boards were measured to permit the fabrication of additional panels beyond those discussed here. Table 2 summarizes the results of the lumber characterization study. The MOE values were adjusted to 12% MC using the procedure defined in ASTM D1990 [24] to allow direct comparison with design values. The average SPFs MOE was significantly higher than expected. The National Design Specification [25] reports the mean MOE for SPFs as 7.58 GPa for No. 2 and 8.27 GPa for No. 1, and the average MOE of the all SPFs used in this study exceeded 8.27 GPa by 34%. While only 4.2% of the SPFs lumber was stamped No. 1, a visual inspection indicated that the vast majority of the SPFs lumber was red spruce (Picea rubens). For comparison, the Wood Handbook [26] gives an average MOE for clear red spruce at 12% MC of 11.45 GPa. In contrast, the LSL MOE was only 1.5% less than the tabulated value of 9.31 GPa [27]. Further, as expected the LSL MOE was much less variable than the SPFs MOE. To ensure that no excessively compliant material was used in CLT panel fabrication, the SPFs boards with MOE values in the lower 5% of the distribution, which corresponded to an MOE of less than 6.89 GPa, were removed from the lot. This shifted the mean MOE to from 11.05 GPa to 11.35 GPa and reduced the coefficient of variation (CoV) in MOE from 19.6% to 15.2%. Following MOE testing, both the SPFs and LSL were conditioned in a dehumidification dry kiln to reduce the MC differential between the two materials and promote better adhesive bonding. The SPFs lumber was conditioned for five days after which it had reached an average MC of 10.8%. The LSL boards were conditioned for 27 days, reaching a MC of 9.4%. The resulting MC differential of 1.4% was well within the recommended moisture content differential of no more than 5% specified in PRG 320 [28].
2.2. Assessment of bond strength The manufacturer-recommended spread rate for the PURBOND adhesive was 100–180 g/m2. This relatively wide range, combined with the uncertainty associated with bonding LSL to SPFs, dictated that an adhesive spread rate study be conducted. To accomplish this, adhesive compression shear block testing was performed per ASTM D905 [29] for adhesive spread rates of 98, 122, 146 and 171 g/m2 for SPFs to SPFs and SPFs to LSL. For each adhesive spread rate and layup, a 127 mm 305 mm two-layer lamination was made from which 10 shear block specimens were cut. Specimens were fabricated from conditioned boards that had been planed to a thickness of 19 mm. Laminates were pressed at 0.01 MPa for two hours per the product standard, and cured per the requirements of ASTM D905. Following each shear block test, strength and percent wood failure were recorded, and each specimen was oven-dried and weighed to determine MC. Results of the shear block tests are given in Table 3. Based on these results, an adhesive spread rate of 146 g/m2 was chosen for CLT panel manufacturing. This adhesive spread rate gave the highest percent wood failure for the SPFs-SPFs specimens, and the highest average shear stress for the SPFs-LSL specimens. We note that the spread rates reported here will likely not be applicable to other brands and types of adhesives.
2.3. Panel fabrication and test specimen preparation Two 2.45 m-long 1.32 m-wide panels of each of the four CLT configurations were fabricated. Both SPFs and LSL boards with minimal warp, twist, bow or cupping were used to ensure reasonable dimensional tolerances. Within two hours of adhesive application, each board was planed to a final thickness of 35 mm, with approximately 1.6 mm removed from each face. Average MC was determined using a pin moisture meter at the time of panel lay-up. Prior to adhesive application, the lumber surface was moistened with a light water spray. A pre-measured amount of PURBOND adhesive was applied using putty knives.
Table 1 CLT Configurations. Configuration
Face Material
Core Material
L1 L2 L3 L4
SPFs (2.8%) LSL LSL SPFs
SPFs LSL SPFs LSL
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Table 2 Summary of results from board E-rating.
SPFs LSL
Avg. MOE (GPa)
Specific Gravity
Density (kg/m3)
MC (%)
11.05 (19.6%) 9.16 (5.1%)
0.42 (2.8%) 0.68 (7.8%)
479 718
14.5 5.0
(CoV in parentheses). (MC = moisture content).
Table 3 Summary of ASTM D905 shear block test results. Lay-Up
Spread Rate (g/m2)
Avg. Shear Stress (MPa)
Average % Wood Failure
Range of % Wood Failure
MC (%)
SPFs SPFs SPFs SPFs SPFs SPFs SPFs SPFs
98 122 146 171 98 122 146 171
11.1 (5.5%) 11.6 (5.9%) 9.4 (8.4%) 10.2 (12.5%) 8.5 (19.5%) 9.0 (20.4%) 10.0 (11.5%) 8.8 (11.7%)
98 90 100 98 75 88 91 98
80–100 40–100 100–100 80–100 50–100 20–100 80–100 95–100
11.9 (1.3%) 11.2 (3.7%) 11.7 (1.5%) 11.9 (2.7%) 6.4 (1.3%) 6.7 (2.9%) 6.4 (2.4%) 6.5 (4.4%)
-
SPFs SPFs SPFs SPFs LSL LSL LSL LSL
(CoV in parentheses). (MC = moisture content).
Following the application of adhesive and panel lay-up, pipe clamps were used to squeeze the panels in the transverse direction and ensure gaps of no more than 3.2 mm between boards in the face layer. The CLT panels were placed in a 1.22 m 2.44 m hydraulic press for two hours at a pressure of 1.03 MPa. With the exception of one all-SPF specimen, which had a gap of 4.8 mm between two adjacent boards in one face layer, the panels met the tolerance requirements of PRG 320 [28]. Three flexure and three shear specimens were cut from each panel. Six specimens were tested in flexure for all four layups, and six specimens were tested in shear for layups L1, L2 and L4.
2.5. Shear test protocol Quasi-static three-point bend tests were conducted in accordance with ASTM D198 [30] to assess shear strength. Load was applied at the middle of the 0.619 m span with a single, 406 mm radius wooden load head. The supports were identical to those employed in the flexural tests. String potentiometers were attached to one side of the specimens at each support and at mid-span (see Fig. 2). As with the flexure tests, each specimen was weighed and its dimensions measured prior to the start of the test, and post-test, oven-dry moisture content was determined. A constant displacement rate of 1.52 mm/min was used, which produced failure after about 10 min.
2.4. Flexure test protocol 2.6. Assessment of specimen stresses at failure from test results Quasi-static four-point flexural tests were conducted in accordance with ASTM D198 [30] as shown in Fig. 1. Load was applied with a single 145 kN hydraulic actuator, and two wooden load heads with a 406 mm radius were attached to a spreader beam. The specimen span was 2.32 m measured between the centerline of the supports, and the load heads contacted the beam at the third points of the simple span. Each support had a steel top plate that rested on a roller allowing free rotation. To ensure that the relatively long supports did not affect the response and that the support reactions were uniformly distributed to the specimen, 102 mm long neoprene bearing pads were sandwiched between the support top plate and the bottom of the beam. Seven string potentiometers were attached at mid-depth on one side of the specimen: one string pot at each support, one at each load head, one midway between each support load head, and one at mid-span. Prior to testing, each specimen was weighed. A pre-load of 1.33 kN was applied, and a load rate of 5 mm/min ensured specimen failure within 6–20 min as specified by ASTM D198 [30]. After loading, failure mode and location were noted, and a 25.4 mm thick cross-sectional slice was taken from an undamaged region of the beam. Dimensions and weight of the cross-sectional slice were recorded, and it was then oven-dried for at least 15 days at 105 °C to determine average moisture content at the time of testing.
Fig. 1. Long-span four-point flexure test.
max
The maximum bending stress f b was determined only for the flexural specimens, and was calculated using Eq. (1). The maximum moment Mmax and effective section modulus Seff were computed using Eqs. (2) and (3), respectively. Pmax is the
Fig. 2. Short-span three-point shear test.
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W.G. Davids et al. / Construction and Building Materials 149 (2017) 156–163 maximum total actuator load, which was evenly split between both load heads. Eq. (2) applies for four-point bending with a span length L and the effective section modulus assumes a transformed section, with the effective depth heff taken as the distance between the extreme fibers of the longitudinal layer(s). PRG 320 [28] recommends that the transformed section modulus Itr be computed assuming a modular ratio n of 30 for the center layer. However, n = 17.6 was used here based on tabulated stiffness properties for red spruce given by Bodig and Jayne [31], since the majority of the lumber used appeared to be red spruce. max
fb
¼
M max Seff
M max ¼
Seff
ð1Þ
ð3Þ max
¼
Pmax Q 2 Ib eff
ð4Þ
P32 Ei h i zi Q ¼ i¼1 EIeff Ib eff
EIeff ¼
3 X
3
Ei bi
i¼1
D0:5L ¼ b
23PL3 1296EIeff
ð11Þ
3. Results and discussion
ð2Þ
2Itr ¼ heff
max
ð10Þ
3.1. Flexure tests: capacities and failure modes
Pmax L 6
The maximum shear stress f v was computed from both the flexural and shear test results using Eqs. (4) and (5) as recommended in the U.S. CLT Handbook [32]. In Eq. (5), the summation upper limit of 3/2 indicates that the shear stress is evaluated at the center of the cross-section. The elastic modulus of the outer layers was taken as the mean value determined from board testing and the elastic modulus of the center layer was taken as E/17.6 for the SPFs. For the LSL, the elastic modulus of the center layer was taken as 1.03 MPa, the value recommended by the LSL manufacturer. The effective elastic bending rigidity EIeff was computed per Eq. (6). The thickness of each layer is hi and zi is the distance from the center of a given layer to the center of the cross-section.
fv
5PL3 324EIeff
Db0:33L ¼
hi þ Ei Ai z2i 12
ð5Þ ! ð6Þ
2.7. Assessment of specimen stiffnesses from flexure test results The load-displacement response for each specimen was used to determine the effective flexural rigidity EIeff , which accounts for only bending deformations; the apparent flexural rigidity EIapp ; which incorporates the effect of both bending and shear deformations; and GAeff ; the effective shear rigidity. All stiffness values were computed based on specimen response between 20% and 50% of maximum measured load Pmax. Therefore, loads generically denoted as P in the following equations are computed as 0:5P max 0:2P max . The term P=D which appears in Eqs. (7)–(9) was computed from a linear regression of the appropriate measured load-displacement response between 0:2P max and 0:5P max . EIeff was computed first for each specimen using Eq. (7), which applies given that shear deformations between the load heads
Typical flexural load-deflection responses for a single specimen of each layup are given in Fig. 3. The reported mid-span deflection is the value measured by the stringpot at mid-span minus the average of the deflections recorded at the support centerlines. Failure load, failure mode, stiffness characteristics and computed stresses are listed for each specimen in Table 4, and averages are presented for each layup in Table 4. The failure loads, stiffness values, and stresses in Table 4 have been adjusted to 12% MC using the methods recommended by ASTM D1990 [24]. Further, reported stresses were calculated based on actual cross-sectional dimensions assuming equal thickness layers and using methods discussed later in this paper. The L2 specimens generally exhibited the lowest degree of variability in both failure load and stiffness, which can be attributed to their being made entirely of LSL, an engineered lumber that typically has more consistent mechanical properties than solid wood. While the failure load CoV was the highest for layup L4, this could be attributed to specimen L4-1 failing at the location of a large knot in the SPFs bottom face layer at a relatively low load of 48.9 kN. Further, all L4 specimens failed in flexural tension of the SPFs bottom layer, and therefore variability in peak load was to a large extent driven by defects in the SPFs. Discarding specimen L4-1 gives a mean failure load of 75.2 kN with CoV of 8.9% for layup L4; this CoV lies between the peak load CoV for lay-up L1 (all SPFs) and L3 (LSL faces and SPFs core). Failure modes were consistent within a given lay-up. All L1 specimens exhibited perpendicular-to-grain shear failures, five of six L2 specimens exhibited flexural tension failures, five of six L3 specimens exhibited flexure tension failures, and all six L4 specimens exhibited flexural tension failures. A typical tension failure is shown in Fig. 4, and a typical shear failure is shown in Fig. 5. The fact that the average bending stress at failure for the L4 spec-
are zero. Here the shear-free deflection within the load span, Dlsb , was computed as the difference between the mid-span displacement and average of the displacements at the load points.
EIeff ¼
P
Dlsb
!
80 L3 432
ð7Þ
EIapp was determined for each specimen using Eq. (8), where Dm is the total midspan displacement minus the average support displacement, and includes both bending and shear deformations.
EIapp ¼
3
P L Dm 56
ð8Þ
Finally, GAeff was computed using Eq. (9), where k was taken as the shear correction factor for a rectangular section of 5/6. Eq. (9) was applied independently at the location of both load heads and mid-span, and the resulting three values of GAeff were averaged to give the reported value for each specimen.
P L Ds 6k
60
Load (kN)
50 40 30 20
ð9Þ
10
In Eq. (9), the shear deflection, Ds , was computed as the total measured deflec-
0
GAeff ¼
tion minus the bending deflection at the load head, D0:33L , and mid-span, D0:5L ; comb b puted with Euler beam theory and the experimentally estimated value of EIeff determined using Eq. (7). Eqs. (10) and (11) apply for computing bending deflections at the load heads and mid-span, respectively.
Specimen L1-1 Specimen L2-3 Specimen L3-3 Specimen L4-3
70
0
10
20
30
40
50
60
Midspan Displacement (mm) Fig. 3. Typical load-displacement response.
70
80
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W.G. Davids et al. / Construction and Building Materials 149 (2017) 156–163
Table 4 Long-span flexure test experimental results. Specimen
Failure Load (kN)
MC (%)
EIeff (kN-m2)
EIapp (kN-m2)
GAeff (kN)
fb
L1-1 L1-2 L1-3 L1-4 L1-5 L1-6 Mean L2-1 L2-2 L2-3 L2-4 L2-5 L2-6 Mean L3-1 L3-2 L3-3 L3-4 L3-5 L3-6 Mean L4-1 L4-2 L4-3 L4-4 L4-5 L4-6 Mean
62.8 54.3 61.5 50.2 61.4 52.2 56.7 48.2 47.8 48.5 46.0 48.5 43.1 47.0 45.9 47.2 45.2 42.3 50.2 52.4 47.2 48.9 80.3 69.5 84.2 70.2 71.6 70.8
11.1 9.2 10.9 9.4 10.1 9.1 10.0 (8.9%) 7.6 7.6 8.9 6.9 8.1 7.5 7.6 (3.8%) 7.6 7.6 8.9 6.9 8.1 7.5 7.8 (8.6%) 9.3 9.9 9.4 9.2 8.5 9.2 9.3 (4.8%)
386 356 340 313 420 392 368 259 285 267 262 265 262 266 264 287 275 275 295 268 277 363 367 317 352 386 322 351
346 297 302 263 322 298 305 236 240 239 230 238 227 235 218 228 231 228 241 236 231 309 336 295 316 336 310 317
6914 3661 5419 3376 2817 2528 4119 5393 3125 4652 3764 4637 3445 4170 2560 2282 2972 2729 2723 3986 2875 5393 3125 4652 3764 4637 3445 7964
44.8 39.3 43.9 35.9 44.1 36.1 40.7 33.9 33.7 34.5 32.7 34.8 30.6 33.4 32.7 33.3 32.3 30.2 35.6 37.4 33.6 34.6 56.8 49.3 59.9 49.3 50.7 50.1
L1
L2
L3
L4
(10.4%)
(4.5%)
(7.7%)
(17.3%)
(10.6%)
(3.5%)
(4.2%)
(7.7%)
(9.1%)
(2.4%)
(3.4%)
(5.2%)
(41.3%)
(20.7%)
(20.5%)
(49.1%)
max
(MPa)
(10.1%)
(4.5%)
(7.5%)
(17.5%)
max
fv
1.40 1.20 1.37 1.12 1.37 1.11 1.26 1.11 1.10 1.11 1.05 1.13 0.98 1.08 1.03 1.07 1.01 0.94 1.13 1.17 1.06 1.12 1.85 1.59 1.91 1.61 1.62 1.62
(MPa)
(10.7%)
(5.0%)
(7.9%)
(17.2%)
Failure Mode Shear Shear Shear Shear Shear Shear – Tension Tension Tension Tension Tension Shear – Tension Tension Tension Tension Tension Shear – Tension Tension Tension Tension Tension Tension –
(CoV in parentheses). max max (MC = moisture content; EIeff = effective bending rigidity; EIapp = apparent bending rigidity; GAeff = effective shear rigidity; f b = maximum bending stress; f v = maximum shear stress).
imens was 23% higher than for the L1 specimens combined with the shift in failure mode from shear to flexural tension supports the premise that incorporating LSL in the core can increase CLT panel shear strength. The L2 and L3 specimens had very similar average failure loads and exhibited consistent flexural tension failures. This indicates that the 1.35E LSL used in the faces of the L2 and L3 panels had a lower average flexural strength than the SPFs used in the faces of the L1 and L4 layups. This may seem counter-intuitive based on the larger allowable flexural stress of 1.35E LSL compared to No. 2 or better SPFs. However, as noted earlier, the lower 5% of the SPFs boards based on MOE were discarded, and relatively straight SPFs boards were selected from the remaining stock for panel manufacturing. This board selection likely increased the overall quality of the SPFs used in panel fabrication.
Fig. 5. Typical shear failure observed in long-span tests.
3.2. Flexure tests: stresses at failure
Fig. 4. Typical flexural tension failure observed in long-span tests.
The nearly identical average bending stress at failure for layups L2 and L3 can be attributed to the fact that for 10 of these 12 specimens, flexural tension failures occurred in the bottom layer of LSL. The highest average flexural stress was for layup L4, where the high shear strength of the LSL core drove flexural tension failures in the bottom SPFs face layer. All of the L1 specimens exhibited max shear failures, which explains why their average f b of 40.7 MPa max is 19% less than the average f b of 50.1 MPa achieved by the L4 max specimens. The average f b of 50.1 MPa for the L4 layup is 9.4 times greater than the allowable bending stress of 5.34 MPa for No. 2 SPFs given by the National Design Specification [25]. While a direct comparison between allowable and observed actual maximum stress should be done with caution, this large discrepancy indicates that the No. 2 or better SPFs used for CLT panel fabrication was of relatively high quality. This is consistent with the high
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W.G. Davids et al. / Construction and Building Materials 149 (2017) 156–163 Table 5 Short-span shear test experimental results. max
Specimen
Failure Load (kN)
MC (%)
fv
L1-1 L1-2 L1-3 L1-4 L1-5 L1-6 Mean L1 L2-1 L2-2 L2-3 L2-4 L2-5 L2-6 Mean L2 L4-1 L4-2 L4-3 L4-4 L4-5 L4-6 Mean L4
79.8 88.1 100.4 95.0 104.8 84.1 92.0 (10.6%) 115.2 117.9 120.4 117.9 106.5 115.4 115.5 (4.2%) 119.0 130.5 109.2 144.0 147.2 142.4 132 (11.6%)
8.3 8.2 8.0 8.3 8.3 7.9 8.2 (2.1%) 6.3 6.5 6.4 6.7 6.2 6.3 6.4 (2.9%) 7.2 6.7 7.6 8.1 7.7 7.7 7.5 (6.2%)
1.76 1.92 2.19 2.11 2.34 1.85 2.03 (10.9%) 2.59 2.68 2.72 2.66 2.41 2.66 2.61 (4.3%) 2.73 2.96 2.51 3.25 3.31 3.04 2.96 (10.3%)
(MPa)
Failure Mode Shear Shear Shear Shear Shear Shear – Tension Shear Tension Tension Tension Shear – Shear Shear Shear Shear Shear Shear –
(CoV in parentheses). max (MC = moisture content; f v = maximum shear stress).
MOE values of the SPFs determined from board testing as discussed earlier. The average shear stress at failure for the L1 lay-up was 1.26 MPa, and should reflect the actual shear strength of the SPFs core material under this loading condition given that all L1 specimax mens failed in shear. This value of f v falls within the shear strength range of 1–2 MPa reported by Sikora et al. [10] for Sitka spruce CLT specimens. Specimen L3-6, the only L3 specimen that failed in shear, had a max computed f v of 1.17 MPa, which is only 7.1% less than average shear stress of 1.26 MPa for the L1 specimens. Computed shear stresses for all other L3 specimens were lower, ranging from 0.94 to 1.13 MPa. The average shear stress in the LSL core of the L4 specimens was 1.62 MPa. If the results for specimen L4-1 are discarded because of its very low failure load driven by a large knot on the bottom face, the average computed shear stress at failure for the remaining five L4 specimens increases to 1.72 MPa. Fig. 6. Typical shear failure observed in short-span tests of L1 specimens.
3.3. Flexure Tests: Stiffnesses
Fig. 7. Typical shear failure observed in short-span tests of L4 specimens.
The L1 lay-up had the highest average EIeff and the L2 lay-up the smallest, which is consistent with the higher average MOE values for the SPFs boards than the LSL boards. The L3 lay-up, with LSL in the faces, had a slightly (4.1%) greater EIeff than the all-LSL L2 lay-up, and EIeff for the L4 lay-up with SPFs in the faces is only 4.6% lower than EIeff for the L1 lay-up. However, when both bending and shear deformations are considered via EIapp ; the L4 lay-up is stiffer than L1. This is likely attributable to the LSL having a greater shear modulus than the SPFs. This is consistent with the larger value of GAeff determined for the L4 lay-up, which reflects the LSL core. The experimentally-determined EIeff values were compared with those determined analytically using the average measured board MOE value of 11.35 GPa for the SPFs, the average measured MOE of 9.16 GPa for the LSL, and transformed section analysis. This approach gives EI values of 320, 259, 258 and 321 kN-m2 for layups L1, L2, L3 and L4, respectively. Comparing with the mean values in Table 4 shows that the analytical approximations compare reasonably well with experimentally determined means. Using
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the experimental values for reference, the errors in approximations are 13%, 3%, 7%, and 9% for layups L1 – L4, respectively. 3.4. Shear tests Based on the flexural test results, short-span shear tests were conducted on layups L1, L2 and L4. Layup L3 was not considered further, since the flexure tests indicate no clear advantage to using low grade LSL face material with an SPFs core. Results for all specimens are provided in Table 5. The L1 specimens all failed in shear (see Fig. 6) at an average max shear stress, f v ; of 2.03 MPa, 61% higher than average shear stress of 1.26 MPa at failure for the L1 flexure specimens. However, bending stresses at failure were 54% higher for the L1 flexure specimens than for the L1 shear specimens. It is possible that the observed L1 flexure specimen shear failures initiated near the load heads and were therefore influenced by simultaneous significant bending stresses. Ig this was the case, the average shear stress at failure of 2.03 MPa may be more representative of the perpendicular-tomax grain shear strength of the SPFs core material. This value of f v is only slightly higher than the upper bound of shear strength of 2 MPa reported by Sikora et al. [10] for Sitka spruce CLT specimens. max The L2 specimens, made entirely of LSL, had an average f v of 2.61 MPa. However, four of the L2 specimens failed in tension, and the shear strength of the LSL therefore cannot be confidently inferred from the L2 data. All the L4 shear specimens exhibited max shear failures (see Fig. 7) with an average f v of 2.96 MPa. This is 83% greater than the average shear stress at failure of 1.62 MPa for the L4 long-span flexure specimens. However, as noted earlier, all the L4 long-span specimens failed in flexural tension, indicating that their average shear stress at failure of 1.62 MPa is a lower bound on LSL shear strength. Therefore, the max shear stress at failure f v of 2.96 MPa is likely the most accurate estimate of the LSL perpendicular-to-grain shear strength, which implies that the LSL has a perpendicular-to-grain shear strength 46% greater than that of the SPFs. Wang et al. [23] reported a mean perpendicular-to-grain shear strength for LSL of only 1.43 MPa, about half the value of 2.96 MPa found in this study. However, Wang et al. [23] also noted that the short-span LSL specimens they tested failed in tension as opposed to shear, which could explain this large difference.
4. Summary and conclusions This study examined the bending and shear performance of three-layer CLT consisting of four-different layups: all SPFs (layup L1), all LSL (lay-up L2), LSL-SPFs-LSL (lay-up L3) and SPFs-LSLSPFs (lay-up L4). Based on long-span bending tests, the L4 panels were the strongest, followed by the L1, L3 and L2 panels. The hybrid L4 panels had a mean bending stress at failure 23% greater than the all-SPFs L1 panels, and the inclusion of low-grade LSL in the core shifted the failure mode from perpendicular-to-grain shear to flexural tension. Short-span shear test results indicate that the low-grade LSL has a shear strength about 46% greater than that of the SPFs. This supports the observation that the inclusion of LSL in the core prevents shear failures in the three-layer CLT panels studied here. These results indicate that hybrid panels with an LSL core may have a structural advantage over all-SPFs panels, allowing for increased span lengths or load-carrying capacity for a given panel span-to-depth ratio. In contrast, the L2 and L3 lay-ups showed only a 3% difference in average bending strength, which reflects the fact that 83% of these panels failed in flexural tension of the LSL face layer. Further, the average bending stress at failure of the all-LSL
L2 specimens was 35% less than that of the L4 specimens, and 16% less than that of the L1 specimens. However, as noted earlier in this paper, a low grade (1.35E) of LSL was used in panel fabrication. Further, based on measured MOE, the SPFs material appears to have been higher quality than typical No. 2 or better material. Therefore, the results reported here may not apply for three-layer CLT panels made from a higher grade of LSL and/or a more typical sample of SPFs material. Further, comparisons between mean strength values do not directly apply to design strengths, which take into account LSL’s inherently low material variability. Finally, the specimens tested here were made from continuous boards as opposed to the finger-jointed lumber typically used in commercially fabricated CLT panels. Finger joints tend to reduce board tensile capacity, and as a result could decrease the differences in flexural strength observed for different layups. Despite these caveats, the inclusion of low-grade LSL as a core material can clearly increase the strength of CLT panels that would otherwise fail in shear.
Acknowledgements This project was supported by Agriculture and Food Research Initiative Competitive Grant no. 2013-34638-21491 from the USDA National Institute of Food and Agriculture. We are also grateful for the materials provided by Pleasant River Lumber, Louisiana-Pacific, and Henkel Adhesives.
References [1] M. Mohammad, S. Gagnon, B.K. Douglas, L. Podesto, Introduction to crosslaminated timber, Wood Des. Focus 22 (2) (2012) 3–12. [2] Y. Chen, F. Lam, Bending performance of box-based cross-laminated timber systems, J. Struct. Eng. 139 (12) (2013) 04013006. [3] A. Robertson, A Comparative Life Cycle Assessment of Mid-rise Office Building Construction Alternatives: Laminated Timber or Reinforced Concrete (Master’s thesis), Univ. of Toronto, Toronto, ON, Canada, 2007. [4] M. Fragiacomo, B. Dujic, I. Sustersic, Elastic and ductile design of multi-storey crosslam massive wooden buildings under seismic actions, Eng. Struct. 33 (11) (2011) 3043–3053. [5] I. Gavric, M. Fragiacomo, A. Ceccotti, Cyclic behavior of CLT wall systems: experimental tests and analytical prediction models, J. Struct. Eng. 141 (11) (2015) 04015034. [6] I. Gavric, M. Fragiacomo, A. Ceccotti, Cyclic behavior of typical metal connectors for cross-laminated (CLT) structures, Mater. Struct. 48 (2015) 1841–1857. [7] S. Pei, J.W. van de Lindt, M. Popovski, Approximate R-factor for cross-laminated timber walls in multistory buildings, J. Archit. Eng. 19 (4) (2013) 245–255. [8] A. Frangi, F. Mario, E. Hugi, R. Jöbst, Experimental analysis of cross-laminated timber panels in fire, Fire Saf. J. 44 (2009) 1078–1087. [9] M. Fragiacomo, A. Menis, I. Clemente, G. Bochicchio, A. Ceccotti, Fire resistance of cross-laminated timber panels loaded out of plane, J. Struct. Eng. 139 (12) (2013) 0401308. [10] K. Sikora, D.O. McPolin, A.M. Harte, Effects of thickness of cross-laminated timber (CLT) panels made from Irish Sitka spruce on mechanical performance in bending and shear, Constr. Build. Mater. 116 (2016) 141–150. [11] D.P. Hindman, J.C. Bouldin, Mechanical properties of southern pine crosslaminated timber, J. Mater. Civ. Eng. 27 (9) (2015) 04014251. [12] H. Ido, H. Nagao, H. Masaki, H. Kato, J. Ogiso, A. Miyatake, Effect of the width and lay-up of sugi cross-laminated timber (CLT) on its dynamic and static elastic moduli, and tensile strength, J. Wood Sci. 62 (2016) 101–108. [13] 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. (2006), http://dx.doi.org/10.1007/s10086-005-0750-7. [14] G. Hochreiner, J. Füssl, J. Eberhardsteiner, Cross-laminated timber plates subjected to concentrated loading: experimental identification of failure mechanisms, Strain 50 (2014) 68–81. [15] G. Hochreiner, J. Füssl, E. Serrano, J. Eberhardsteiner, Influence of wooden board strength class on the performance of cross-laminated timber plates investigated by means of full-field deformation measurements, Strain 50 (2) (2014) 161–173. [16] J. Vessby, B. Enquist, H. Petersson, T. Alsmarker, Experimental study of crosslaminated timber wall panels, Eur. J. Wood Prod. 67 (2009) 211–218. [17] T. Bogensperger, T. Moosbrugger, G. Silly, Verification of CLT-plates under loads in plane. Proceedings of the 11th World Conference on Timber Engineering, Riva del Garda, Italy, 2010.
W.G. Davids et al. / Construction and Building Materials 149 (2017) 156–163 [18] 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. [19] Q. Zhou, M. Gong, Y.H. Chiu, M. Mohammad, Measurement of rolling shear modulus and strength of cross laminated timber fabricated with black spruce, Constr. Build. Mater. 64 (2014) 379–386. [20] M. Li, F. Lam, Y. Li. Evaluating rolling shear strength properties of crosslaminated timber by torsional shear tests, in: Proceedings of the 2014 World Conference on Timber Engineering, Quebec City, Canada, August 10–14, 2014. [21] Y. Li, F. Lam, Low cycle fatigue tests and damage accumulation models of the rolling shear strength of cross-laminated timber, J. Wood Sci. 62 (2016) 251– 262. [22] S. Aicher, Z. Christian, M. Hirsch, Rolling shear modulus and strength of beech wood laminations, Holzforschung (2016), http://dx.doi.org/10.1515/hf-20150229. [23] Z. Wang, M. Gong, Y.-H. Chui, Mechanical properties of laminated strand lumber and hybrid cross-laminated timber, Constr. Build. Mater. 101 (2016) 622–627. [24] ASTM, D1990-14 Standard Practice for Establishing Allowable Properties for Visually-Graded Dimension Lumber From in-Grade Tests of Full-Size Specimens, American Society of Testing and Materials, West Conshohocken, PA, 2014. ASTM International.
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[25] American Wood Council, National Design Specification: Design Values for Wood Construction, 2012 ed., American Wood Council, Leesburg, VA, 2012. [26] Forest Products Laboratory, Wood Handbook - Wood as an Engineering Material, U.S. Department of Agriculture, Madison, WI, 2010. [27] ICC Evaluation Service, LP SOLIDSTART Laminated Strand Lumber (LSL) and Laminated Veneer Lumber (LVL). Technical Report No. ESR-2403, ICC Evaluation Service, 2015. [28] American National Standard Institute & APA – The Engineered Wood Association, ANSI/APA PRG 320–2012, in Standard for Performance-rated Cross-Laminated Timber. Tacoma, WA: APA – The Engineered Wood Association, 2012. [29] ASTM, D905-08 Standard Test Method for Strength Properties of Adhesive Bonds in Shear by Compression Loading, 2013 ed., American Society of Testing and Materials, West Conshohocken, PA, 2008. ASTM International. [30] ASTM, D198-14 standard test methods of static testing of lumber in structural sizes, in: American Society of Testing And Materials, ASTM International, West Conshohocken, PA, 2014. [31] J. Bodig, B.A. Jayne, Mechanics of Wood and Wood Composites, Van Nostrand Reinhold Company, New York, NY, 1982. [32] FPInnovations & Binational Softwood Lumber Council. CLT Handbook: CrossLaminated Timber (E. Karacabeyli & B. Douglas, Eds.). Pointe-Claire, QC: FPInnovations, 2013.
Contents lists available at ScienceDirect
Construction and Building Materials journal homepage: www.elsevier.com/locate/conbuildmat
Structural performance of hybrid SPFs-LSL cross-laminated timber panels William G. Davids a,⇑, Nicholas Willey a, Roberto Lopez-Anido a, Stephen Shaler b, Douglas Gardner b, Russell Edgar c, Mehdi Tajvidi b a b c
Department of Civil and Environmental Engineering, University of Maine, 5711 Boardman Hall, Orono, MA 04469-5711, USA School of Forest Resources, University of Maine, 5755 Nutting Hall, Orono, MA 04469-5755, USA Advanced Structures and Composites Center, University of Maine, 35 Flagstaff Road, Orono, MA 04469-5793, USA
h i g h l i g h t s Hybrid CLT with laminated strand lumber (LSL) tested under out of plane loads. LSL core eliminated shear failure in 3-layer panels. CLT panel bending strength increased by 23% through inclusion of LSL core.
a r t i c l e
i n f o
Article history: Received 6 December 2016 Received in revised form 10 May 2017 Accepted 13 May 2017
Keywords: Cross-laminated timber Laminated strand lumber Wood shear strength Wood flexural strength Structural testing
a b s t r a c t The bending and shear performance of hybrid cross-laminated timber (CLT) panels made from SprucePine-Fir (South) (SPFs) and laminated strand lumber (LSL) are examined. Four configurations of threelayer CLT were fabricated: all-SPFs control specimens, all-LSL specimens, hybrid specimens with SPFs faces and an LSL core, and hybrid specimens with LSL faces and an SPFs core. Bending tests were conducted to assess flexural strength and stiffness. Additionally, three-point bending tests were performed to assess shear performance. The incorporation of LSL in the core of CLT panels increased mean panel bending stress at failure by 23% through mitigation of rolling shear failure. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Originally introduced in Austria and Germany in the mid-1990s, cross-laminated timber (CLT) has become an increasingly popular alternative for multi-story timber construction in Europe [1]. CLT has recently garnered interest in North America with the establishment of several CLT and nail-laminated timber plants in Canada and the United States. CLT panels are suitable for use in walls, floors and roofs, and are typically fabricated from an odd number of flat-wise layers of solid-sawn lumber placed in alternating 90 degree directions. In the majority of cases, individual layers of boards are adhesively bonded although nail- and screwlaminated CLT is also produced. Alternative forms of CLT have been considered including placing laminations at ±45 degrees as well as hollow, box-based systems [2]. Compared to typical concrete con⇑ Corresponding author. E-mail address: [email protected] (W.G. Davids). http://dx.doi.org/10.1016/j.conbuildmat.2017.05.131 0950-0618/Ó 2017 Elsevier Ltd. All rights reserved.
struction, CLT structures are lightweight, sequester more carbon, possess better thermal insulation properties, and are more rapidly erected [3]. Research on the structural performance of CLT can be separated into the broad categories of seismic behavior [4–7], fire resistance [8,9], and determination of the mechanical properties of CLT. A number of studies have focused on the determination of CLT mechanical properties in flexure and shear, which are primary design properties for panels subjected to out-of-plane loading [10–15]. Sikora et al. [10] present a current and thorough review of the existing literature on this topic. Others have focused on CLT mechanical response due to in-plane loading [16–18]. As with plywood, an issue which can limit the capacity of CLT subjected to out-of-plane loading is failure in perpendicular-tograin shear, commonly called rolling shear. Rolling shear also contributes to deflections of CLT panels. Because of its significance, several investigations have considered rolling shear properties and failure mechanisms. Zhou et al. [19] examined the effect of
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rolling shear deformations in 3-layer, black spruce CLT, measuring rolling shear modulus and conducting three-point bending tests of CLT specimens. Zhou et al. [19] also proposed a deflection adjustment factor to account for rolling shear deformations, and also concluded that bending specimen width did not significantly affect apparent elastic modulus and apparent shear modulus. Li et al. [20] implemented a torsional test for evaluating rolling shear strength in CLT, observing that thinner cross-layers tended to have higher rolling shear strengths. Hochreiner et al. [14] studied CLT plates subjected to concentrated loads and examined the evolution of rolling shear failure modes by tracking fracture development and load-deformation history using digital image correlation. Li and Lam [21] experimentally assessed rolling shear damage accumulation in CLT attributable to load cycling, and calibrated a damage accumulation model that can be used for future studies on duration-of-load behavior of CLT under rolling shear. Noting the significance of rolling shear on CLT structural performance, Aicher et al. [22] assessed the rolling shear modulus and strength of European beech, which typically has much better rolling shear properties than softwoods normally used in CLT construction, concluding that the use of beech in CLT cross-layers could be beneficial for CLT strength. Wang et al. [23] assessed the use of laminated strand lumber (LSL) in both cross-layers and face layers of hybrid CLT panels, demonstrating increased flexural capacities relative to conventional all-softwood CLT. The literature indicates that rolling shear failure can be a limiting factor for the strength of CLT subjected to out-of-plane loading. The focus of the research reported in this paper was the structural assessment of hybrid CLT panels made from LSL and softwood lumber with the objective of increasing strength by mitigating rolling shear failures in the core layer. LSL is an engineered composite lumber that is made from approximately 300 mm long strands of fast-growing species (often aspen or poplar) that are bonded and densified during manufacture and oriented with the long axis of the structural member. LSL typically possesses good dimensional stability and very predictable strength and stiffness values compared to solid-sawn lumber. Additionally, the authors are aware of no published experimental research specifically examining the use of Northeastern U.S. Spruce-Pine-Fir (South) (SPFs) lumber in CLT. SPFs is an economically significant group of lumber species harvested in the United States that includes Eastern Spruces, Balsam Fir, Red Pine, Jack Pine, Englemann Spruce, Lodgepole Pine, Sitka Spruce and Norway Spruce. All species are subject to the same grading rules and have the same design values. CLT panel production using SPFs harvested and milled in the Northeastern US may become an important new market for lumber producers in the United States as CLT markets grow. The research reported in this paper includes characterization of the SPFs and LSL lumber used for CLT manufacturing, assessment of the bond between SPFs and LSL using a polyurethane adhesive, and testing to assess both major-axis flexural and shear strength and stiffness.
2. Materials and methods Materials were 38 mm 184 mm 3 m kiln-dried SPFs No. 2, 38 mm 184 mm 3 m grade 1.35E LSL boards (without wax coating on the board edges), and Henkel PURBOND HB E452 single-component polyurethane adhesive. The No. 2 grade of SPFs is a standard grading category corresponding to specific stiffness and strength design values in the US codes, and is a commonly produced grade of SPFs lumber. The SPFs lumber was procured in bulk quantities from Pleasant River Lumber in Dover-Foxcroft, Maine, USA, and as discussed later, a small percentage of the SPFs was No. 1, a higher grade with higher design values. The LSL was provided by Louisiana-Pacific Corporation’s plant in Houlton, Maine, USA. The designation ‘‘1.35E LSL” refers to a specific grade of LSL that has a nominal elastic modulus of 9310 MPa. The 1.35E LSL was selected as opposed to a higher grade – 1.55E and 1.75E grades with moduli of 10,700 MPA and 12,070 MPA are also available – because its flexural strength and stiffness were expected to be similar to the SPFs.
Two, three-layer CLT panels were laid up for each of four configurations as described in Table 1. The CLT panels were made from continuous boards as opposed to the finger-jointed lumber typically used in commercially fabricated CLT panels.
2.1. Lumber characterization and preparation Each SPFs and LSL board was first E-rated using a Metriguard 340 E-Computer, its moisture content (MC) taken with a Delhorst J2000 pin moisture meter, and its dimensions measured and density calculated. In the E-rating process, the board is placed flatwise on two supports, one of which contains a small load cell. The board is struck at mid-span with a hammer, and the dynamic load cell readings are used to compute a dynamic modulus of elasticity (MOE). A total of over 900 SPFs and over 700 LSL boards were measured to permit the fabrication of additional panels beyond those discussed here. Table 2 summarizes the results of the lumber characterization study. The MOE values were adjusted to 12% MC using the procedure defined in ASTM D1990 [24] to allow direct comparison with design values. The average SPFs MOE was significantly higher than expected. The National Design Specification [25] reports the mean MOE for SPFs as 7.58 GPa for No. 2 and 8.27 GPa for No. 1, and the average MOE of the all SPFs used in this study exceeded 8.27 GPa by 34%. While only 4.2% of the SPFs lumber was stamped No. 1, a visual inspection indicated that the vast majority of the SPFs lumber was red spruce (Picea rubens). For comparison, the Wood Handbook [26] gives an average MOE for clear red spruce at 12% MC of 11.45 GPa. In contrast, the LSL MOE was only 1.5% less than the tabulated value of 9.31 GPa [27]. Further, as expected the LSL MOE was much less variable than the SPFs MOE. To ensure that no excessively compliant material was used in CLT panel fabrication, the SPFs boards with MOE values in the lower 5% of the distribution, which corresponded to an MOE of less than 6.89 GPa, were removed from the lot. This shifted the mean MOE to from 11.05 GPa to 11.35 GPa and reduced the coefficient of variation (CoV) in MOE from 19.6% to 15.2%. Following MOE testing, both the SPFs and LSL were conditioned in a dehumidification dry kiln to reduce the MC differential between the two materials and promote better adhesive bonding. The SPFs lumber was conditioned for five days after which it had reached an average MC of 10.8%. The LSL boards were conditioned for 27 days, reaching a MC of 9.4%. The resulting MC differential of 1.4% was well within the recommended moisture content differential of no more than 5% specified in PRG 320 [28].
2.2. Assessment of bond strength The manufacturer-recommended spread rate for the PURBOND adhesive was 100–180 g/m2. This relatively wide range, combined with the uncertainty associated with bonding LSL to SPFs, dictated that an adhesive spread rate study be conducted. To accomplish this, adhesive compression shear block testing was performed per ASTM D905 [29] for adhesive spread rates of 98, 122, 146 and 171 g/m2 for SPFs to SPFs and SPFs to LSL. For each adhesive spread rate and layup, a 127 mm 305 mm two-layer lamination was made from which 10 shear block specimens were cut. Specimens were fabricated from conditioned boards that had been planed to a thickness of 19 mm. Laminates were pressed at 0.01 MPa for two hours per the product standard, and cured per the requirements of ASTM D905. Following each shear block test, strength and percent wood failure were recorded, and each specimen was oven-dried and weighed to determine MC. Results of the shear block tests are given in Table 3. Based on these results, an adhesive spread rate of 146 g/m2 was chosen for CLT panel manufacturing. This adhesive spread rate gave the highest percent wood failure for the SPFs-SPFs specimens, and the highest average shear stress for the SPFs-LSL specimens. We note that the spread rates reported here will likely not be applicable to other brands and types of adhesives.
2.3. Panel fabrication and test specimen preparation Two 2.45 m-long 1.32 m-wide panels of each of the four CLT configurations were fabricated. Both SPFs and LSL boards with minimal warp, twist, bow or cupping were used to ensure reasonable dimensional tolerances. Within two hours of adhesive application, each board was planed to a final thickness of 35 mm, with approximately 1.6 mm removed from each face. Average MC was determined using a pin moisture meter at the time of panel lay-up. Prior to adhesive application, the lumber surface was moistened with a light water spray. A pre-measured amount of PURBOND adhesive was applied using putty knives.
Table 1 CLT Configurations. Configuration
Face Material
Core Material
L1 L2 L3 L4
SPFs (2.8%) LSL LSL SPFs
SPFs LSL SPFs LSL
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Table 2 Summary of results from board E-rating.
SPFs LSL
Avg. MOE (GPa)
Specific Gravity
Density (kg/m3)
MC (%)
11.05 (19.6%) 9.16 (5.1%)
0.42 (2.8%) 0.68 (7.8%)
479 718
14.5 5.0
(CoV in parentheses). (MC = moisture content).
Table 3 Summary of ASTM D905 shear block test results. Lay-Up
Spread Rate (g/m2)
Avg. Shear Stress (MPa)
Average % Wood Failure
Range of % Wood Failure
MC (%)
SPFs SPFs SPFs SPFs SPFs SPFs SPFs SPFs
98 122 146 171 98 122 146 171
11.1 (5.5%) 11.6 (5.9%) 9.4 (8.4%) 10.2 (12.5%) 8.5 (19.5%) 9.0 (20.4%) 10.0 (11.5%) 8.8 (11.7%)
98 90 100 98 75 88 91 98
80–100 40–100 100–100 80–100 50–100 20–100 80–100 95–100
11.9 (1.3%) 11.2 (3.7%) 11.7 (1.5%) 11.9 (2.7%) 6.4 (1.3%) 6.7 (2.9%) 6.4 (2.4%) 6.5 (4.4%)
-
SPFs SPFs SPFs SPFs LSL LSL LSL LSL
(CoV in parentheses). (MC = moisture content).
Following the application of adhesive and panel lay-up, pipe clamps were used to squeeze the panels in the transverse direction and ensure gaps of no more than 3.2 mm between boards in the face layer. The CLT panels were placed in a 1.22 m 2.44 m hydraulic press for two hours at a pressure of 1.03 MPa. With the exception of one all-SPF specimen, which had a gap of 4.8 mm between two adjacent boards in one face layer, the panels met the tolerance requirements of PRG 320 [28]. Three flexure and three shear specimens were cut from each panel. Six specimens were tested in flexure for all four layups, and six specimens were tested in shear for layups L1, L2 and L4.
2.5. Shear test protocol Quasi-static three-point bend tests were conducted in accordance with ASTM D198 [30] to assess shear strength. Load was applied at the middle of the 0.619 m span with a single, 406 mm radius wooden load head. The supports were identical to those employed in the flexural tests. String potentiometers were attached to one side of the specimens at each support and at mid-span (see Fig. 2). As with the flexure tests, each specimen was weighed and its dimensions measured prior to the start of the test, and post-test, oven-dry moisture content was determined. A constant displacement rate of 1.52 mm/min was used, which produced failure after about 10 min.
2.4. Flexure test protocol 2.6. Assessment of specimen stresses at failure from test results Quasi-static four-point flexural tests were conducted in accordance with ASTM D198 [30] as shown in Fig. 1. Load was applied with a single 145 kN hydraulic actuator, and two wooden load heads with a 406 mm radius were attached to a spreader beam. The specimen span was 2.32 m measured between the centerline of the supports, and the load heads contacted the beam at the third points of the simple span. Each support had a steel top plate that rested on a roller allowing free rotation. To ensure that the relatively long supports did not affect the response and that the support reactions were uniformly distributed to the specimen, 102 mm long neoprene bearing pads were sandwiched between the support top plate and the bottom of the beam. Seven string potentiometers were attached at mid-depth on one side of the specimen: one string pot at each support, one at each load head, one midway between each support load head, and one at mid-span. Prior to testing, each specimen was weighed. A pre-load of 1.33 kN was applied, and a load rate of 5 mm/min ensured specimen failure within 6–20 min as specified by ASTM D198 [30]. After loading, failure mode and location were noted, and a 25.4 mm thick cross-sectional slice was taken from an undamaged region of the beam. Dimensions and weight of the cross-sectional slice were recorded, and it was then oven-dried for at least 15 days at 105 °C to determine average moisture content at the time of testing.
Fig. 1. Long-span four-point flexure test.
max
The maximum bending stress f b was determined only for the flexural specimens, and was calculated using Eq. (1). The maximum moment Mmax and effective section modulus Seff were computed using Eqs. (2) and (3), respectively. Pmax is the
Fig. 2. Short-span three-point shear test.
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W.G. Davids et al. / Construction and Building Materials 149 (2017) 156–163 maximum total actuator load, which was evenly split between both load heads. Eq. (2) applies for four-point bending with a span length L and the effective section modulus assumes a transformed section, with the effective depth heff taken as the distance between the extreme fibers of the longitudinal layer(s). PRG 320 [28] recommends that the transformed section modulus Itr be computed assuming a modular ratio n of 30 for the center layer. However, n = 17.6 was used here based on tabulated stiffness properties for red spruce given by Bodig and Jayne [31], since the majority of the lumber used appeared to be red spruce. max
fb
¼
M max Seff
M max ¼
Seff
ð1Þ
ð3Þ max
¼
Pmax Q 2 Ib eff
ð4Þ
P32 Ei h i zi Q ¼ i¼1 EIeff Ib eff
EIeff ¼
3 X
3
Ei bi
i¼1
D0:5L ¼ b
23PL3 1296EIeff
ð11Þ
3. Results and discussion
ð2Þ
2Itr ¼ heff
max
ð10Þ
3.1. Flexure tests: capacities and failure modes
Pmax L 6
The maximum shear stress f v was computed from both the flexural and shear test results using Eqs. (4) and (5) as recommended in the U.S. CLT Handbook [32]. In Eq. (5), the summation upper limit of 3/2 indicates that the shear stress is evaluated at the center of the cross-section. The elastic modulus of the outer layers was taken as the mean value determined from board testing and the elastic modulus of the center layer was taken as E/17.6 for the SPFs. For the LSL, the elastic modulus of the center layer was taken as 1.03 MPa, the value recommended by the LSL manufacturer. The effective elastic bending rigidity EIeff was computed per Eq. (6). The thickness of each layer is hi and zi is the distance from the center of a given layer to the center of the cross-section.
fv
5PL3 324EIeff
Db0:33L ¼
hi þ Ei Ai z2i 12
ð5Þ ! ð6Þ
2.7. Assessment of specimen stiffnesses from flexure test results The load-displacement response for each specimen was used to determine the effective flexural rigidity EIeff , which accounts for only bending deformations; the apparent flexural rigidity EIapp ; which incorporates the effect of both bending and shear deformations; and GAeff ; the effective shear rigidity. All stiffness values were computed based on specimen response between 20% and 50% of maximum measured load Pmax. Therefore, loads generically denoted as P in the following equations are computed as 0:5P max 0:2P max . The term P=D which appears in Eqs. (7)–(9) was computed from a linear regression of the appropriate measured load-displacement response between 0:2P max and 0:5P max . EIeff was computed first for each specimen using Eq. (7), which applies given that shear deformations between the load heads
Typical flexural load-deflection responses for a single specimen of each layup are given in Fig. 3. The reported mid-span deflection is the value measured by the stringpot at mid-span minus the average of the deflections recorded at the support centerlines. Failure load, failure mode, stiffness characteristics and computed stresses are listed for each specimen in Table 4, and averages are presented for each layup in Table 4. The failure loads, stiffness values, and stresses in Table 4 have been adjusted to 12% MC using the methods recommended by ASTM D1990 [24]. Further, reported stresses were calculated based on actual cross-sectional dimensions assuming equal thickness layers and using methods discussed later in this paper. The L2 specimens generally exhibited the lowest degree of variability in both failure load and stiffness, which can be attributed to their being made entirely of LSL, an engineered lumber that typically has more consistent mechanical properties than solid wood. While the failure load CoV was the highest for layup L4, this could be attributed to specimen L4-1 failing at the location of a large knot in the SPFs bottom face layer at a relatively low load of 48.9 kN. Further, all L4 specimens failed in flexural tension of the SPFs bottom layer, and therefore variability in peak load was to a large extent driven by defects in the SPFs. Discarding specimen L4-1 gives a mean failure load of 75.2 kN with CoV of 8.9% for layup L4; this CoV lies between the peak load CoV for lay-up L1 (all SPFs) and L3 (LSL faces and SPFs core). Failure modes were consistent within a given lay-up. All L1 specimens exhibited perpendicular-to-grain shear failures, five of six L2 specimens exhibited flexural tension failures, five of six L3 specimens exhibited flexure tension failures, and all six L4 specimens exhibited flexural tension failures. A typical tension failure is shown in Fig. 4, and a typical shear failure is shown in Fig. 5. The fact that the average bending stress at failure for the L4 spec-
are zero. Here the shear-free deflection within the load span, Dlsb , was computed as the difference between the mid-span displacement and average of the displacements at the load points.
EIeff ¼
P
Dlsb
!
80 L3 432
ð7Þ
EIapp was determined for each specimen using Eq. (8), where Dm is the total midspan displacement minus the average support displacement, and includes both bending and shear deformations.
EIapp ¼
3
P L Dm 56
ð8Þ
Finally, GAeff was computed using Eq. (9), where k was taken as the shear correction factor for a rectangular section of 5/6. Eq. (9) was applied independently at the location of both load heads and mid-span, and the resulting three values of GAeff were averaged to give the reported value for each specimen.
P L Ds 6k
60
Load (kN)
50 40 30 20
ð9Þ
10
In Eq. (9), the shear deflection, Ds , was computed as the total measured deflec-
0
GAeff ¼
tion minus the bending deflection at the load head, D0:33L , and mid-span, D0:5L ; comb b puted with Euler beam theory and the experimentally estimated value of EIeff determined using Eq. (7). Eqs. (10) and (11) apply for computing bending deflections at the load heads and mid-span, respectively.
Specimen L1-1 Specimen L2-3 Specimen L3-3 Specimen L4-3
70
0
10
20
30
40
50
60
Midspan Displacement (mm) Fig. 3. Typical load-displacement response.
70
80
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Table 4 Long-span flexure test experimental results. Specimen
Failure Load (kN)
MC (%)
EIeff (kN-m2)
EIapp (kN-m2)
GAeff (kN)
fb
L1-1 L1-2 L1-3 L1-4 L1-5 L1-6 Mean L2-1 L2-2 L2-3 L2-4 L2-5 L2-6 Mean L3-1 L3-2 L3-3 L3-4 L3-5 L3-6 Mean L4-1 L4-2 L4-3 L4-4 L4-5 L4-6 Mean
62.8 54.3 61.5 50.2 61.4 52.2 56.7 48.2 47.8 48.5 46.0 48.5 43.1 47.0 45.9 47.2 45.2 42.3 50.2 52.4 47.2 48.9 80.3 69.5 84.2 70.2 71.6 70.8
11.1 9.2 10.9 9.4 10.1 9.1 10.0 (8.9%) 7.6 7.6 8.9 6.9 8.1 7.5 7.6 (3.8%) 7.6 7.6 8.9 6.9 8.1 7.5 7.8 (8.6%) 9.3 9.9 9.4 9.2 8.5 9.2 9.3 (4.8%)
386 356 340 313 420 392 368 259 285 267 262 265 262 266 264 287 275 275 295 268 277 363 367 317 352 386 322 351
346 297 302 263 322 298 305 236 240 239 230 238 227 235 218 228 231 228 241 236 231 309 336 295 316 336 310 317
6914 3661 5419 3376 2817 2528 4119 5393 3125 4652 3764 4637 3445 4170 2560 2282 2972 2729 2723 3986 2875 5393 3125 4652 3764 4637 3445 7964
44.8 39.3 43.9 35.9 44.1 36.1 40.7 33.9 33.7 34.5 32.7 34.8 30.6 33.4 32.7 33.3 32.3 30.2 35.6 37.4 33.6 34.6 56.8 49.3 59.9 49.3 50.7 50.1
L1
L2
L3
L4
(10.4%)
(4.5%)
(7.7%)
(17.3%)
(10.6%)
(3.5%)
(4.2%)
(7.7%)
(9.1%)
(2.4%)
(3.4%)
(5.2%)
(41.3%)
(20.7%)
(20.5%)
(49.1%)
max
(MPa)
(10.1%)
(4.5%)
(7.5%)
(17.5%)
max
fv
1.40 1.20 1.37 1.12 1.37 1.11 1.26 1.11 1.10 1.11 1.05 1.13 0.98 1.08 1.03 1.07 1.01 0.94 1.13 1.17 1.06 1.12 1.85 1.59 1.91 1.61 1.62 1.62
(MPa)
(10.7%)
(5.0%)
(7.9%)
(17.2%)
Failure Mode Shear Shear Shear Shear Shear Shear – Tension Tension Tension Tension Tension Shear – Tension Tension Tension Tension Tension Shear – Tension Tension Tension Tension Tension Tension –
(CoV in parentheses). max max (MC = moisture content; EIeff = effective bending rigidity; EIapp = apparent bending rigidity; GAeff = effective shear rigidity; f b = maximum bending stress; f v = maximum shear stress).
imens was 23% higher than for the L1 specimens combined with the shift in failure mode from shear to flexural tension supports the premise that incorporating LSL in the core can increase CLT panel shear strength. The L2 and L3 specimens had very similar average failure loads and exhibited consistent flexural tension failures. This indicates that the 1.35E LSL used in the faces of the L2 and L3 panels had a lower average flexural strength than the SPFs used in the faces of the L1 and L4 layups. This may seem counter-intuitive based on the larger allowable flexural stress of 1.35E LSL compared to No. 2 or better SPFs. However, as noted earlier, the lower 5% of the SPFs boards based on MOE were discarded, and relatively straight SPFs boards were selected from the remaining stock for panel manufacturing. This board selection likely increased the overall quality of the SPFs used in panel fabrication.
Fig. 5. Typical shear failure observed in long-span tests.
3.2. Flexure tests: stresses at failure
Fig. 4. Typical flexural tension failure observed in long-span tests.
The nearly identical average bending stress at failure for layups L2 and L3 can be attributed to the fact that for 10 of these 12 specimens, flexural tension failures occurred in the bottom layer of LSL. The highest average flexural stress was for layup L4, where the high shear strength of the LSL core drove flexural tension failures in the bottom SPFs face layer. All of the L1 specimens exhibited max shear failures, which explains why their average f b of 40.7 MPa max is 19% less than the average f b of 50.1 MPa achieved by the L4 max specimens. The average f b of 50.1 MPa for the L4 layup is 9.4 times greater than the allowable bending stress of 5.34 MPa for No. 2 SPFs given by the National Design Specification [25]. While a direct comparison between allowable and observed actual maximum stress should be done with caution, this large discrepancy indicates that the No. 2 or better SPFs used for CLT panel fabrication was of relatively high quality. This is consistent with the high
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W.G. Davids et al. / Construction and Building Materials 149 (2017) 156–163 Table 5 Short-span shear test experimental results. max
Specimen
Failure Load (kN)
MC (%)
fv
L1-1 L1-2 L1-3 L1-4 L1-5 L1-6 Mean L1 L2-1 L2-2 L2-3 L2-4 L2-5 L2-6 Mean L2 L4-1 L4-2 L4-3 L4-4 L4-5 L4-6 Mean L4
79.8 88.1 100.4 95.0 104.8 84.1 92.0 (10.6%) 115.2 117.9 120.4 117.9 106.5 115.4 115.5 (4.2%) 119.0 130.5 109.2 144.0 147.2 142.4 132 (11.6%)
8.3 8.2 8.0 8.3 8.3 7.9 8.2 (2.1%) 6.3 6.5 6.4 6.7 6.2 6.3 6.4 (2.9%) 7.2 6.7 7.6 8.1 7.7 7.7 7.5 (6.2%)
1.76 1.92 2.19 2.11 2.34 1.85 2.03 (10.9%) 2.59 2.68 2.72 2.66 2.41 2.66 2.61 (4.3%) 2.73 2.96 2.51 3.25 3.31 3.04 2.96 (10.3%)
(MPa)
Failure Mode Shear Shear Shear Shear Shear Shear – Tension Shear Tension Tension Tension Shear – Shear Shear Shear Shear Shear Shear –
(CoV in parentheses). max (MC = moisture content; f v = maximum shear stress).
MOE values of the SPFs determined from board testing as discussed earlier. The average shear stress at failure for the L1 lay-up was 1.26 MPa, and should reflect the actual shear strength of the SPFs core material under this loading condition given that all L1 specimax mens failed in shear. This value of f v falls within the shear strength range of 1–2 MPa reported by Sikora et al. [10] for Sitka spruce CLT specimens. Specimen L3-6, the only L3 specimen that failed in shear, had a max computed f v of 1.17 MPa, which is only 7.1% less than average shear stress of 1.26 MPa for the L1 specimens. Computed shear stresses for all other L3 specimens were lower, ranging from 0.94 to 1.13 MPa. The average shear stress in the LSL core of the L4 specimens was 1.62 MPa. If the results for specimen L4-1 are discarded because of its very low failure load driven by a large knot on the bottom face, the average computed shear stress at failure for the remaining five L4 specimens increases to 1.72 MPa. Fig. 6. Typical shear failure observed in short-span tests of L1 specimens.
3.3. Flexure Tests: Stiffnesses
Fig. 7. Typical shear failure observed in short-span tests of L4 specimens.
The L1 lay-up had the highest average EIeff and the L2 lay-up the smallest, which is consistent with the higher average MOE values for the SPFs boards than the LSL boards. The L3 lay-up, with LSL in the faces, had a slightly (4.1%) greater EIeff than the all-LSL L2 lay-up, and EIeff for the L4 lay-up with SPFs in the faces is only 4.6% lower than EIeff for the L1 lay-up. However, when both bending and shear deformations are considered via EIapp ; the L4 lay-up is stiffer than L1. This is likely attributable to the LSL having a greater shear modulus than the SPFs. This is consistent with the larger value of GAeff determined for the L4 lay-up, which reflects the LSL core. The experimentally-determined EIeff values were compared with those determined analytically using the average measured board MOE value of 11.35 GPa for the SPFs, the average measured MOE of 9.16 GPa for the LSL, and transformed section analysis. This approach gives EI values of 320, 259, 258 and 321 kN-m2 for layups L1, L2, L3 and L4, respectively. Comparing with the mean values in Table 4 shows that the analytical approximations compare reasonably well with experimentally determined means. Using
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the experimental values for reference, the errors in approximations are 13%, 3%, 7%, and 9% for layups L1 – L4, respectively. 3.4. Shear tests Based on the flexural test results, short-span shear tests were conducted on layups L1, L2 and L4. Layup L3 was not considered further, since the flexure tests indicate no clear advantage to using low grade LSL face material with an SPFs core. Results for all specimens are provided in Table 5. The L1 specimens all failed in shear (see Fig. 6) at an average max shear stress, f v ; of 2.03 MPa, 61% higher than average shear stress of 1.26 MPa at failure for the L1 flexure specimens. However, bending stresses at failure were 54% higher for the L1 flexure specimens than for the L1 shear specimens. It is possible that the observed L1 flexure specimen shear failures initiated near the load heads and were therefore influenced by simultaneous significant bending stresses. Ig this was the case, the average shear stress at failure of 2.03 MPa may be more representative of the perpendicular-tomax grain shear strength of the SPFs core material. This value of f v is only slightly higher than the upper bound of shear strength of 2 MPa reported by Sikora et al. [10] for Sitka spruce CLT specimens. max The L2 specimens, made entirely of LSL, had an average f v of 2.61 MPa. However, four of the L2 specimens failed in tension, and the shear strength of the LSL therefore cannot be confidently inferred from the L2 data. All the L4 shear specimens exhibited max shear failures (see Fig. 7) with an average f v of 2.96 MPa. This is 83% greater than the average shear stress at failure of 1.62 MPa for the L4 long-span flexure specimens. However, as noted earlier, all the L4 long-span specimens failed in flexural tension, indicating that their average shear stress at failure of 1.62 MPa is a lower bound on LSL shear strength. Therefore, the max shear stress at failure f v of 2.96 MPa is likely the most accurate estimate of the LSL perpendicular-to-grain shear strength, which implies that the LSL has a perpendicular-to-grain shear strength 46% greater than that of the SPFs. Wang et al. [23] reported a mean perpendicular-to-grain shear strength for LSL of only 1.43 MPa, about half the value of 2.96 MPa found in this study. However, Wang et al. [23] also noted that the short-span LSL specimens they tested failed in tension as opposed to shear, which could explain this large difference.
4. Summary and conclusions This study examined the bending and shear performance of three-layer CLT consisting of four-different layups: all SPFs (layup L1), all LSL (lay-up L2), LSL-SPFs-LSL (lay-up L3) and SPFs-LSLSPFs (lay-up L4). Based on long-span bending tests, the L4 panels were the strongest, followed by the L1, L3 and L2 panels. The hybrid L4 panels had a mean bending stress at failure 23% greater than the all-SPFs L1 panels, and the inclusion of low-grade LSL in the core shifted the failure mode from perpendicular-to-grain shear to flexural tension. Short-span shear test results indicate that the low-grade LSL has a shear strength about 46% greater than that of the SPFs. This supports the observation that the inclusion of LSL in the core prevents shear failures in the three-layer CLT panels studied here. These results indicate that hybrid panels with an LSL core may have a structural advantage over all-SPFs panels, allowing for increased span lengths or load-carrying capacity for a given panel span-to-depth ratio. In contrast, the L2 and L3 lay-ups showed only a 3% difference in average bending strength, which reflects the fact that 83% of these panels failed in flexural tension of the LSL face layer. Further, the average bending stress at failure of the all-LSL
L2 specimens was 35% less than that of the L4 specimens, and 16% less than that of the L1 specimens. However, as noted earlier in this paper, a low grade (1.35E) of LSL was used in panel fabrication. Further, based on measured MOE, the SPFs material appears to have been higher quality than typical No. 2 or better material. Therefore, the results reported here may not apply for three-layer CLT panels made from a higher grade of LSL and/or a more typical sample of SPFs material. Further, comparisons between mean strength values do not directly apply to design strengths, which take into account LSL’s inherently low material variability. Finally, the specimens tested here were made from continuous boards as opposed to the finger-jointed lumber typically used in commercially fabricated CLT panels. Finger joints tend to reduce board tensile capacity, and as a result could decrease the differences in flexural strength observed for different layups. Despite these caveats, the inclusion of low-grade LSL as a core material can clearly increase the strength of CLT panels that would otherwise fail in shear.
Acknowledgements This project was supported by Agriculture and Food Research Initiative Competitive Grant no. 2013-34638-21491 from the USDA National Institute of Food and Agriculture. We are also grateful for the materials provided by Pleasant River Lumber, Louisiana-Pacific, and Henkel Adhesives.
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