KEYWORDS: Cross-laminated timber, fire resistance, model, tests

Transcription

1 FIRE-EXPOSED CROSS-LAMINATED TIMBER MODELLING AND TESTS Joachim Schmid 1, Jürgen König, Jochen Köhler 3 ABSTRACT: Cross-laminated timber (CLT) is increasingly used in medium-rise timber buildings, for the time being up to eight storeys, among other reasons for cost effectiveness and robustness. This paper presents a simple design model using the effective cross-section method for the structural fire design, i.e. the determination of the mechanical resistance with respect to bending (floors). Performing advanced calculations for a large number of lay-ups of various lamination thicknesses, using the thermal and thermo-mechanical properties of wood, charring depths and the reduction of bending resistance of CLT were determined as functions of time of fire exposure. From these results zero-strength layers were derived to be used in the design model using an effective residual cross-section for the determination of mechanical resistance. The model also takes into account different temperature gradients in the CLT in order to include the effect of slower heating rate when the CLT is protected by insulation and/or gypsum plasterboard. The paper also gives results from fire-tests of CLT in bending using beam strips cut from CLT with adequate side protection in order to achieve one-dimensional heat transfer. Reference tests at ambient temperature were performed to predict the moment resistance of the beams being tested in fire. KEYWORDS: Cross-laminated timber, fire resistance, model, tests 1 INTRODUCTION 13 Cross-laminated timber (CLT) made of flatwise laminations are increasingly used as structural elements in housing and commercial buildings, both in floors and walls. Recently eight storeys high buildings were erected in Sweden. The laminations are held together by bonding. In structural fire design, for timber beams and columns EN [1] gives a simplified method for the calculation of the mechanical resistance (reduced cross-section model). Apart from charring that reduces the size of the cross-section of the member, the effect of elevated temperature in a 35 to mm deep zone below the char layer is taken into account by assuming a zerostrength layer of depth 7 mm immediately below the char layer, and ambient strength and stiffness properties in the remaining effective residual cross-section. For unprotected timber members, this zero-strength layer is assumed to increase linearly from zero to its maximum value during the first minutes of the fire exposure. When EN was drafted, it was believed that the reduced cross-section method could be applied to timber slabs exposed to fire on one side using a zero-strength layer of 7 mm. At that time the use of CLT in housing 1 Joachim Schmid, SP Trätek, Box 569, SE Stockholm, Sweden. joachim.schmid@sp.se Jürgen König, SP Trätek, Box 569, SE Stockholm, Sweden. juergen.koenig@sp.se 3 Jochen Köhler, Swiss Federal Institute of Technology ETH, Zurich, Switzerland. jochen.koehler@ibk.baug.ethz.ch was just in the beginning. It was not intended that the reduced cross-section method could be applied to this new type of construction without further investigation. In a preliminary study [] on timber decks in bending, it was shown that the application of a zero-strength layer of 7 mm would give unsafe results in many applications. This paper presents an easy-to-use design model for CLT decks (floors) and a comparison with the results from fire tests. CLT walls will be subject of future publications. COMPUTER SIMULATIONS.1 ASSUMPTIONS.1.1 Thermal analysis The thermal analyses were performed using SAFIR 7 [3] using thermal properties of wood given by EN [1]. The thermal properties of gypsum plasterboard were similar to those given in [], however further calibrated to fit test results, see [5]..1. Structural analysis For the structural analysis, a computer program CSTFire, written as a Visual Basic macro embedded in Excel, was developed, using the temperature output from the heat transfer calculations and the relative strength and stiffness values given by EN [1], i.e. compressive and tensile strengths, f c and f t, and moduli of elasticity in compression and tension, E c and E t. These values are given as bi-linear functions of temperature from to 3 C with breakpoints at 1 C, also taking

2 into account the effects of transient moisture situations and creep, see Figure 1. The software takes into account the possibility of permitting ductile behaviour of wood under elevated temperature. Contrary to ambient conditions where failure on the tension side of a beam is brittle, in the fire situation tensile failure of the outermost fibres won t cause immediate collapse of the member since a redistribution of internal stresses will take place as long as equilibrium is maintained. Reduction factor k Θ 1,5 f v f c E c f t E t 1 3 Temperature [ C] Figure 1: Reduction factors for strength and stiffness properties according to EN [1]. σ f t f c ε C 6 C 1 C C Figure : Temperature-dependent stress-strain relationships for wood for different temperatures. Since the reduction of strength and stiffness properties is different for tension and compression, CSTFire uses an iteration process, increasing the curvature of the member until the maximum moment resistance is reached. The element size used for the thermal and structural analysis was chosen as 1 mm 1 mm. The calculations were conducted assuming material properties that are representative for timber deck plates used in practice, using the stress-strain relationship shown in Figure. Since the values of tensile and compressive strength of solid timber given in design or product standards, e.g. EN 338 [6], are values related to the whole cross-section and were determined on the assumption of a linear relationship between stress and strain until failure, the use of these values in a finite element analysis would not be correct []. Therefore, compressive strength values were determined using the data from Thunell [7] dependent on density and moisture content as shown in []. For the timber slabs assumed here the compressive strength was f c = 3 N/mm, while the tensile strength was taken as f t = 7 N/mm, that is the ratio of f t /f c is,9. The results of simulations of crosslaminated timber slabs were taken from [5]. These calculations were made assuming f t /f c = (,9). Since the ratio has only a small influence on the results, it can be neglected for practical application. For layers with the grain direction perpendicular to the longitudinal direction of the plate (cross-layers), the modulus of elasticity was assumed to be zero, while these layers were assumed to be completely effective with respect to shear stiffness, i.e. complete composite action of the longitudinal layers was assumed. This assumption is justified by the slenderness ratios of CLT occurring in practice. For deck plates of thickness 15 mm and a span of 5,5 m, the slenderness ratio is 3. Due to shear deformations the maximum bending stress is increased by about 3 % [8]. Since the slenderness ratio in fire due to charring is somewhat greater than under ambient conditions, the increase of bending stresses due to shear deformations would be less than 3 %. This influence is neglected in the model presented here. The calculations were carried out for members in bending with depths from 5 to 315 mm, layer thicknesses from 15 to 5 mm and layer numbers from 3 to 7, both for unprotected and protected CLT, and both for the fire-exposed side in tension or compression. Both regular and irregular lay-ups were studied, i.e. regular lay-ups with equal layer thicknesses of longitudinal and cross-layers, and irregular ones where layer thicknesses of longitudinal and cross-layers were different. In all cases, however, all longitudinal layers had the same thickness and the lay-ups were symmetrical.. Results In the following, for unprotected CLT with five layers of mm thickness, the results of the computer simulations are shown as relationships between the bending moment ratio M fi /M and time t and zero-strength layer d and time, respectively, see Figure and Figure 5 (for CLT with three or seven layers, see [5]. In one case, the relationships are also shown as functions of the charring depth d char (Figure 6). The zero-strength layer d, see Figure 3, is used by EN [1] for the reduced cross-section method in order to take into account the reduction of strength and stiffness properties due to elevated temperatures.

3 n = i n = i-1 n = 3 n = n = 1 h hef d dchar n = i n = i-1 n = 3 n = n = 1 Figure 3: Effective residual cross-section obtained after reduction of original cross-section (left) with charring depth d char and zero-strength layer d (right). d [mm] , d EN d,fi,fm,rescs EN d [mm] t [min] d d = 7mm, (d = 7 mm) Figure : Results for CLT with 5 layers of thickness mm with the fire-exposed side in tension. The yellow bars indicate charring of bondline. The black broken bars indicate the and % levels of bending resistance. d [mm] t [min] d d = 7 mm, EN Figure 5: Results for CLT with 5 layers of thickness mm with the fire-exposed side in compression. The yellow bars indicate charring of bondline. The black broken bars indicate the and % levels of bending resistance. Figure 6: Same results as in Figure 5, however shown as functions of charring depth d char. For the examples shown here, during the first charring phase when charring takes place in the first (loadbearing) layer, the bending resistance is reduced to approximately forty percent of ambient resistance. In the beginning of the next charring phase when charring takes place in the second (non load-bearing) layer, the decay of bending resistance is very slow, but increasing considerably before the char front has reached the second bond line. At that stage the bending resistance has dropped to twenty-five to twenty percent, depending on the state of stress on the fire-exposed side. When the fire-exposed side is in compression, the reduction of bending resistance is greatest. The charts also show the corresponding zero-strength layers d that should be applied to the cross-section to get the same results using ambient strength and stiffness properties for the effective residual cross-section. Since the value of d varies considerable with time (or charring depth d char ), in order to simplify the design model, the largest value within the bending resistance interval between and percent was chosen. In Figure and Figure 5 these values are marked as rings. For the CLT plates shown, d = 1,5 mm for the fire-exposed side in tension and d = 15,9 mm for the fire-exposed side in compression. EN [1] gives, for rectangular cross-sections (beams and columns), a general value for the zerostrength layer as d = 7 mm, once a stabilized temperature profile has established after about twenty minutes. From the charts can be seen that, for the most relevant stage of relative resistances from to, this value would give non-conservative results in comparison with the results from the simulations. This calculation was carried out for a large number of lay-ups with five layers. The data for d where the plotted as functions of the depth h of the CLT plate, see Figure 7. The trendlines, somewhat modified and simplified, are given as: For unprotected CLT with five layers where the fireexposed side is in tension

4 h d = + 1 in millimetres (1) 1 For unprotected CLT with five layers where the fireexposed side is in compression h d = + 11 in millimetres () d [mm] layers h [mm] Exposed side in compression Exposed side in tension d [mm] d = 7mm d (tf = 69 min) d,model d (no protection failure (tf = 69 min) (d = 7 mm) (no protection failure), Figure 7: Determination of linear expression for zerostrength layer d vs. plate depth for unprotected CLT. For protected CLT, the relationships of M fi /M and d are shown as functions of time t and charring depth d char, respectively, see Figure 8 and Figure 9. d [mm] t [min] d = 7mm d d (model) d (tf = 69 min) (no protection failure) (tf = 69 min) (d = 7mm), Figure 8: Results for CLT with 5 layers of thickness mm with the fire-exposed side in compression and one layer of 15 mm gypsum plasterboard protection. The yellow bars indicate charring of bondline. The black broken bars indicate the and % levels of bending resistance. Figure 9: Same results as in Figure 8, however shown as functions of charring depth d char. The protection chosen in the simulations was a 15, mm thick layer of gypsum plasterboard. Two cases were assumed: 1. The gypsum plasterboard remains in place,. The gypsum plasterboard falls off after 69 minutes (using a failure criterion of 8 C being reached on the unexposed side of the gypsum plasterboard; it is just a coincidence that the relative resistance is at that time). There is a large effect of the protection on the fire resistance, c.f. Figure 5 and Figure 8 since the start of charring is delayed and the charring rate behind the gypsum plasterboard is slower than in the unprotected case. Comparing Figure 6 with Figure 9, however, we can see that, for the same charring depth (which are valid for different times t), the bending resistance is smaller when the CLT is protected. Since the heating rate is smaller due to the protection provided by the gypsum plasterboard the temperature gradient in the wood is smaller than in the unprotected case. Therefore, failure of the gypsum plasterboard at time t f has a positive effect on the bending resistance (and the zero-strength layer d ), however the total effect of failure of the gypsum plasterboard is negative, since the charring rate increases considerably during the post-protection phase. The simulations for protected CLT were carried out for a large number of regular lay-ups and protective claddings consisting of one layer (1,5 mm or 15, mm thick) or two layers 1,5 + 15, mm thick gypsum plasterboard. The gypsum plasterboards were assumed to fall off when the temperature in the interface between gypsum plasterboard and wood was 7,, 6 and 8 C respectively. The largest value of d within the relative

5 resistance interval between and % was determined and plotted versus the depth h of the CLT, similar to Figure 7. The expressions found are given by: For protected CLT with the fire-exposed side in tension h d = 3 for 75 mm h 1 mm (3) h d = + 6 for h > 1 mm () 35 For protected CLT with the fire-exposed side in compression d = 18 mm 3 TESTS 3.1 GENERAL In order to verify the model by tests, two CLT products currently being on the market were chosen. The width of all specimens was 15 mm. The specimens of series M had a lay-up of five layers of equal thickness 19 mm (regular lay-up), i.e. the total thickness was 95 mm, while the lay-up of series S with a thickness of 15 mm was irregular: (in millimetres), see Figure 1. The width and length of the specimens chosen was governed by the test conditions in the fire situation. Both are considerably smaller than required for testing of CLT products. EN 789 [9] requires a span of 3 times the depth of the CLT panel, while the minimum width required by the standard is 3 mm. A sample of each product was subdivided into two groups, one of which was tested in fire (series MF and SF) while the other was tested at ambient conditions (series MR and SR) in order to provide data for the prediction of the ambient bending moment resistance of the specimens to be tested in fire. In CLT panels lamellae finger joints are situated randomly. Since finger joints in a lamellae of series SF could have a considerable influence on the bending resistance in fire [1], the specimens of series SR and SF were produced or selected such that they did not contain any finger joints in the most stressed parts of the beams. Since the lamellae of CLT in series M, in practice normally used for walls, are butt-jointed; i.e. the buttjoints are non-loadbearing in tension), the test specimens were produced without any joints in the longitudinal layers. This product will therefore exhibit lower bending strength when the CLT contains butt-joints. All test specimens were conditioned at C and 65 % relative humidity. Series M (MR and MF) Series S (SR and SF) Figure 1: Lay-ups of specimens in series M and S, in millimetres. 3. REFERENCE TESTS UNDER AMBIENT CONDITIONS The ambient reference tests were carried out as fourpoint ramp load tests. For series SR the span was,7 m and the two point loads were acting at the third points (,9 m +,9 m +,9 m). For series MR, in order to reduce the risk of shear failure, the distances were chosen to be equal as in the fire tests (,9 m + 1,5 m +,9 m, see Figure 1). The number of tests was 1 (series MR) and 15 (series SR), respectively. Two different failure modes were observed, either tensile failure in the outer lamella or shear failure in one or two of the cross-layers see Figure 11 and Figure 1. The results are shown in Figure 13. In series MR one out of 1 specimens failed due to shear failure while the rest failed due to tensile failure. In Series SR seven specimens failed due to tensile failure and eight specimens due to shear failure. Since shear failure was not expected in the fire tests and is not a relevant failure mode in the simulations and the design model, the results of series SR and MR were evaluated with respect to the relevant failure mode failure of tensile lamella. The parameters of a lognormal distribution have been estimated by using the censored using the Maximum Likelihood Method [11]. Figure 11: Example of tensile failure in reference test of series SR.

6 Figure 1: Example of shear failure in reference test of series SR. By using the censored Maximum Likelihood Method, the different failure modes and their logical arrangement can be considered. In the case of a serial arrangement only the realisation of the lowest strength can be detected directly. For a sample of n observations of system failure it can by judgement be distinguished between the two failure modes. Each failure mode is observed an arbitrary number of times. It is assumed that the strength according to failure mode tension of the outermost lamella, which can be modelled by the random variable X, is observed k times (x 1, x,, x k ). The parameters of the probability distribution function of X can be found considering (x 1, x,, x k ) and the observed strength values (s 1, s,, s n-k ) where the other failure mechanism ( shear ) is involved. Two likelihood functions are formulated, the first one considering (x 1, x,, x k ) quantitatively: k L = f 1 i= 1 ( x θ) i (11)( (5) The second likelihood function uses the information that n k strength values s i are smaller than a realisation of X: L n k = i=1 P ( X s θ) i (1) (6) With P X s θ = 1 F s θ (13) (7) ( ) ( ) i i The optimal set of distribution parameters can be found by solving the maximization problem: θ ˆ max L L (8) = θ ( ) 1 The results are shown in Figure 13. The moments of distribution functions are given in Table 1: Moments of the distributions for tension of the outermost lamella. MR SR Mean Value [N/mm ] 58,7 5 Standard Deviation [N/mm ] 5,1 1, 1,9,7,5,3, SR Bending Failure MR Bending Failure SR LogNormal f m [N/mm ] SR Shear Failure MR Shear Failure MR LogNormal Figure 13: Results from ambient reference tests for series MR and SR plotted together with the corresponding probability distributions (LogNormal, parameters obtained with censored maximum likelihood). 3.3 FIRE TESTS General and test procedure Fire tests of specimens in bending were performed for the following cases: Unprotected timber; fire-exposed side in tension (tsw); Unprotected timber; fire-exposed side in compression (csw); Protected timber; fire-exposed side in tension (tsw); Protected timber; fire-exposed side in compression (csw). In order to simulate the thermal conditions in a CLT plate in fire, i.e. the one-dimensional heat flux, and consequently one dimensional charring, the edges of the beam specimens were protected by a first layer of mm thick pieces of wood with the grain in longitudinal direction and a second layer of 15 mm thick pieces of gypsum plasterboard type F, all of them fixed with nails. These layers were discontinuous in order to prevent composite action, with gap widths of 1 mm. Where the CLT beams were protected, 15 mm long pieces of gypsum plasterboard type F were screwed to the bottom face of the beams with 1 mm gaps between the pieces. The supports of the specimens were located on the furnace walls outside the heated zone of one metre, see Figure 1. The loads could be applied in upward or downward direction. A measuring device was placed on top of the specimens for measuring the deflection within a gauge length of 9 mm.

7 Figure 15: Examples of residual cross-sections after fire test. The photographs taken of the residual cross-sections were used to record the borderline of the shape using the software AutoCAD. From these data, the area, second moment of area and the section modulus were determined. Since the charring depth was not equal over the width of the beams, the area was used for the calculation of the mean charring depth for a crosssection. For the determination of the second moment of area, only load-bearing layers (with the grain in longitudinal direction) were considered. mm Figure 1: Test furnace, location of test specimen and loading equipment [1]. The load was applied prior to the fire test, then the furnace was started and the load was kept constant until failure. Unlike the failure modes observed during the ambient reference tests, failure in fire was preceded by extensive deflections. At bending failure or when the load could not be held constant, the burners were turned off, the specimen removed from the furnace and the fire in the wood extinguished with water. The time elapsed from turning off the burners to extinguishing the fire was normally from 1 to 1,5 minutes Test results From each of the test specimens, at five locations, photographs were taken of the cross-section after the fire test, see two examples shown in Figure 15. Specimen SF 11 (left) was unprotected while specimen SF 1 (right) was protected. Especially for protected beams it was difficult to achieve one-dimensional heat flux and charring over the whole width of the beams, due to opening of wide gaps between the bottom and side protection. Comparisons of test results with the simulations using the dimensions of the test beams are shown in Figure 16 to Figure 3 with relationships of the bending resistance ratio versus charring depth. For each test, the bending moment resistance was predicted using the results from the reference tests at ambient temperature. The graphs also show the relative bending resistance obtained using the simplified model for the zero-strength layer, and the value that would be obtained applying a zero-strength layer of d = 7 mm as given in EN [1]. Since the model was fitted to give best results in the range of a relative bending resistances between and, the values are only shown for this interval. In general, the test results agree fairly well with the simulations. Some deviations are, however, noticeable. A more in-depth analysis of the specimens after fire tests showed that some specimens exhibited local deficiencies of charring depth, caused by char ablation or other. This may have caused lower bending resistances of some tests shown in Figure 18. Since such local deficiencies are more effective in narrow beams, it can be argued that CLT of sizes used in practice are less vulnerable to local defects. The simplified model for the zero-strength layer normally gives conservative results compared with the simulations, except for protected CLT. Non-conservative deviations are due to the assumption that the cladding would fall off after some time; during the fire tests, however, the protective cladding remained in place during the fire tests. The assumption of a zero-strength layer equal 7 mm gives unsafe results, especially when the fire-exposed side is in compression.

8 Series MF Unprotected tsw Series SF Unprotected csw, Simulation Model d = 7 mm Figure 16: Comparison of Test results with simulation and simplified design model for series MF, unprotected, with the fire-exposed side in tension (tsw)., Simulation Series MF Unprotected csw Model d = 7 mm Figure 17: Comparison of Test results with simulation and simplified design model for series MF, unprotected, with the fire-exposed side in compression (csw)., Simulation Model d = 7 mm Figure 19: Comparison of Test results with simulation and simplified design model for series SF, unprotected, with the fire-exposed side in compression (csw)., Series MF Protected tsw Simulation Model d = 7 mm Figure : Comparison of Test results with simulation and simplified design model for series MF, protected, with the fire-exposed side in tension (tsw). Series SF Unprotected tsw Series MF Protected csw, Simulation Model d = 7 mm Figure 18: Comparison of Test results with simulation and simplified design model for series SF, unprotected, with the fire-exposed side in tension (tsw)., Simulation Model d = 7 mm Figure 1: Comparison of Test results with simulation and simplified design model for series MF, protected, with the fire-exposed side in compression (csw).

9 , Simulation Series SF Protected tsw Model d = 7 mm Figure : Comparison of Test results with simulation and simplified design model for series SF, protected, with the fire-exposed side in tension (tsw)., Simulation Series SF Protected csw Model d = 7 mm Figure 3: Comparison of Test results with simulation and simplified design model for series SF, protected, with the fire-exposed side in compression (csw). CONCLUSIONS It has been shown that the complex performance of CLT exposed to fire can be described by advanced computer simulations, using the thermal and thermo-mechanical properties of wood given by EN [1] and that the simulation results are verified by test results from fire tests. In order to present a user-friendly easy-to-use design model for members in bending, the concept of the reduced cross-section method given in [1] was adopted and zero-strength layers determined for consideration of the reduced strength and stiffness properties at elevated temperatures. The simplified model gives reliable results, while the adoption of the zero-strength layer equal to 7 mm, as given in EN [1] for beams and columns, normally gives non-conservative results. ACKNOWLEDGEMENT The research described here was conducted at SP Trätek, Stockholm, as a part of the FireInTimber project within the European Wood-Wisdom-Net framework. It is supported by industry through the European Initiative Building With Wood and public funding organisations. The test specimens were produced and delivered by Martinsons Trä (Sweden) and Stora Enso Austria. A part of the fire tests were assisted, evaluated and reported by Per Willinder to be included in his Bachelor thesis [1]. REFERENCES [1] EN Eurocode 5: Design of timber structures Part 1-: General Structural fire design. European Committee for Standardization, Brussels,. [] König, J, Schmid, J. Bonded timber deck plates in fire. Proceedings of CIB W18, Meeting, Bled, Slovenia. Lehrstuhl für Ingenieurholzbau, University of Karlsruhe, Karlsruhe, Germany, 7. [3] Franssen, J.M, Kodur, V.K.R., Mason, J., User s manual for SAFIR A computer program for analysis of structures subjected to fire. University of Liege, Department Structures du Génie Civil Service Ponts et Charpentes. Liege, Belgium, 5. [] Källsner, B. and König, J., Thermal and mechanical properties of timber and some other materials used in light timber frame construction. Proceedings of CIB W18, Meeting 33, Delft, Lehrstuhl für Ingenieurbau, University Karlsruhe, Karlsruhe, Germany,. [5] Schmid, J, König, J. Cross-laminated timber in fire. Research report SP, in preparation, to be published 1. [6] EN 338:3, Structural timber Strength classes. European Standard. European Committee for Standardization, Brussels, 3. [7] Thunell, B., Hållfasthetsegenskaper hos svenskt furuvirke utan kvistar och defekter. Royal Swedish Institute for Engineering Research, Proceedings No. 161, Stockholm, 191. [8] Blass, H, Görlacher, R, Brettsperrholz Grundlagen. Holzbau-Kalender. 3 () S [9] EN 789:, Timber structures Test methods Determination of mechanical properties of wood based panels. European Standard. European Committee for Standardization, Brussels,. [1] König, J, Norén, J, Sterley, M, Effect of adhesives on finger joint performance in fire. CIB W18, Meeting 1, St. Andrews, Canada. Lehrstuhl für Ingenieurholzbau, University of Karlsruhe, Karlsruhe, Germany, 8. [11] Steiger, R, Köhler, J. Analysis of censored data examples in timber engineering research. Proceedings of CIB W18, Meeting 38, Karlsruhe, Lehrstuhl für Ingenieurbau, University Karlsruhe, Karlsruhe, Germany, 5. [1] Willinder, P. Fire resistance of cross-laminated timber. Bachelor thesis. Jönköping University, Jönköping, Sweden, 1.

10