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Lightweight and high-performance concrete
Many researchers studied the influence of high-performance concrete on the performance of timber-concrete composite (TCC). In general, high-performance concrete, such as lightweight concrete, steel fiber reinforced concrete, high strength concrete, and self-compacting concrete, was deemed more advantageous than standard concrete. The timber design with such concretes shows some edges in terms of mass, stiffness, resistance, and thickness but comes with a higher cost. Steinberg et al. [49] used lightweight concrete to minimize the total mass of the structures. It was concluded that the structures were affected by the low modulus of elasticity of lightweight concrete, and this causes the reduction of the effective bending stiffness. A denser connector layout could be used to compensate for this reduction. The lightweight concrete did not affect the stiffness and resistance of the proprietary connector “Tecnaria” [21] since the timber properties governed the connector performance. However, a higher grade lightweight concrete was also recommended [50] to assure connection efficiency.
Steel fiber reinforced concrete was used in the study of Kieslich and Holschemacher [51] to reduce the thickness of the concrete slab. Results from the push-out test showed that shear strength and initial stiffness are increased compared to regular reinforced concrete. Lamothe et al. [52] stated the same conclusion based on the experimental tests of the TCC notched connector using ultra-high performance fiber reinforced concrete. The notch connector was more rigid, more robust, and helped avoid shear failure of concrete. Using lightweight concrete could reduce floor mass. The steel fiber reinforced concrete is a potential solution to minimize floor thickness while retaining the same performance, consequently limiting the floor mass. However, the floor mass reduction could lead to inferior vibration performance based on Ghafar’s study [44] about the impact of concrete thickness on the natural frequencies.
For notch connectors, shrinkage of concrete in the early days of curing duration will result in a gap at the outer edge of the connection. The phenomenon caused undesired initial permanent deflection of the composite beam. Therefore, it is recommended to use low shrinkage concrete to minimize this unwanted phenomenon while obtaining high workability and flowability [53], [54].
Notch reinforced with fastener
Notch connector is a prominent solution for TCC structure using mass timber panel. Notches cut could be made by a simple CNC machining process during the fabrication of timber panels. In addition, the notch with vertical screws could be a remedy to the low ductility of the notch connector [25]. However, the ductility of individual connectors depends on the type and the number of screws, and a ductile connector could not ensure the ductile behavior of a whole structure [43].
In the study of Rijal [61], it was concluded that the bird-mouth type connections exhibited higher strength and stiffness than the trapezoidal notch connections. The failure mode of bird-mouth notched connections was different from trapezoidal notched connections. The latter had to crush and splitting failure in the LVL and bending of the lag screw, while the former had no damage in the notch and no bending of the lag screw. This difference in the failure mode could probably be due to the high tensile strength of the lag screw used in these series.
A combination of notch and lag screw was investigated by Deam et al. [58]. A lag screw could increase the performances of the specimens with the round concrete notch. The rectangular notch with the lag screw provided greater strength and stiffness with a larger notch area. The failure mode of these two types of a combination is reported to be the same as the lag screw appeared to hold the fractured surfaces of the concrete together. The strength dropped when the lag screw eventually fractured.
Gutkowski et al. [62] tested the shear performance of many configurations of notch size with Hilti dowel and Borden resin (Figure 1.11). The result shows that the notch dimension affected the slip modulus and strength of 2×4 specimens and made no significant effect on 4×4 specimens. Various failure modes were observed in the slip test, with none being predominant.
Human perception toward floor vibration
The human perception of floor vibration is complex and challenging to measure. The sensitivity and subjectivity of the human body lead to the fact that no limit is stated for acceptable vibration levels in the design of the building, but only guidelines were developed. The sensitivity of human perception on the floor vibration can be evaluated by the acceleration and the velocity responses to the fundamental natural frequency. Many correlations of subjective perception to an easy-to-use design guideline were found in the literature and discussed in Section 0. These guidelines focus on different vibration responses such as fundamental frequency, number of natural frequencies below 40 Hz, damping, mean acceleration, peak acceleration, velocity, and deflection under a specified static load.
Negreira et al. [70] conducted an extensive psycho-vibratory evaluation of timber floors in laboratory conditions using multilevel regression. The authors demonstrated the relationship between the subjective answers of the floor occupants and many measured vibration responses of the floors. The results showed that the best indicator for vibration annoyance is the fundamental frequency (calculated based on EC5 guidelines [8]) and Hu and Chui’s ratio [71] (calculated using the fundamental frequency and the deflection of the floor under 1kN point load). On the other hand, for vibration acceptance, the best indicator is the Maximum Transient Vibration Value (a computed based on the acceleration experienced by the test subjects, as per Standard ISO 2631-1:1997 [72]).
Experimental studies on CCC structures in literature
Many researchers conducted experimental tests on the CCC flooring structures to determine the static and dynamic properties. Before conducting investigations on CCC beams, an assessment of composite connectors was usually performed.
For example, Mai et al. [93] performed dynamic and static tests on 6-m CCC beams based on an experimental investigation on a screw composite connector [4]. Higgins et al. [5] test CCC beams using screw connectors at Oregon State University. Lamothe et al. [52] tested 9m beams of CLT-HPC (High-performance concrete) composite. The authors used a bird-mouth notched connector with reinforced screws. Zhang et al. [45], [94] focused on the influence of geometry on the notched connector of CCC floors. Jiang and Crocetti [6] tested the notched connector’s shear properties and the CCC beam’s bending resistance using this type of connector.
The static aspects of CCC beams were studied extensively recently, and the results showed that this construction system could robustly withstand short- and long-term loading. On the other hand, CCC beam dynamic behavior was less studied even though vibration performance usually governs the CCC design, especially for the long-span beams and floors.
Shear tests on connectors
The performance characteristics of connectors for serviceability and ultimate limit state (SLS and ULS) can be determined through the direct shear push-out test according to standard EN 26891:1991 [95]. The strength is quantified as the maximum shear load 𝐹𝑚𝑎𝑥 applied when the failure occurs in the push-out specimen and defined as the highest value of shear force monitored during the test for slips not larger than 15 mm. Maximum shear resistance 𝐹𝑚𝑎𝑥 is estimated before conducting the test. The stiffness is quantified by the slip modulus at three different load levels (40, 60, and 80% of the mean maximum load) corresponding to the service, ultimate, and near-collapse load levels [60] (Figure 1.14). As the vibrational problems lie in the serviceability limit, the behavior of the structure is considered linear elastic. Therefore, the vibrational performance of the composite timber-concrete structure highly depends on the stiffness, or more specifically, the stiffness at 40 % of estimated failure load, of the connection systems.
Influence of heel length
Figure 2.6 exhibits the relationship between the heel length and the variables of interest, namely K1, K2, and Fmax. One can see an increase of about 15% of the stiffness K1 and K2 when the heel length increases from 300 mm to 400 mm (Figure 2.6.a and b). Heel length was assumed not to influence either the stiffness or strength of the connector (Figure 2.6.c). This slight increase was probably because of the asymmetrical properties of the test. The lengthy heel magnified the eccentricity and the friction between concrete and timber. The resistance of the connector of different notch depths distinguished clearly from each other’s; they developed almost independently regarding their heel length. Modulus K2 was more consistent than K1 since the specimen was stabilized after the first loading sequence. In terms of the effect of heel length on the failure type, a specimen with a shorter heel tended to have its lamellas sheared off at failure. The error bars in the graphs represent the 95% confidence interval of the mean value 𝑥̅. They are calculated as 𝑥̅±𝑡𝑛−1.𝑠/√𝑛, with 𝑠 is the standard deviation of the sample, 𝑛 is the sample size, and 𝑡𝑛−1 is the upper (1−0.95)/2 critical value for the t distribution with (n−1) degrees of freedom. Since the standard error was significant in some average data points, the evolution of K1 and K2 was challenging to be verified.
Influence of notch depth
Figure 2.7.a and b show the evolution of slip modulus when the notch depth increases. A deeper cut did not yield a stiffer connector. The notch with 25 mm depth had the highest stiffness in most cases. In the notch with 35 mm of depth, the timber material of the first layer of the lamella was extracted entirely, and the second layer, which laid in the direction perpendicular to the first one, was weaker in terms of modulus perpendicular to the grain. The transversal timber lamellas were also not glued edgewise. They could be the reason for the “peak” trend of the slip modulus curves. Figure 2.7.a and b show the stiffness K1 and K2 per notch depth. The modulus gained per millimeter of notch depth was higher in the less deep notches. The shallow notch used the material more effectively in terms of stiffness, and further extraction of material in the topmost longitudinal layer would reduce the effectiveness of the connector. The linear correlations between the stiffness per depth and the notch depth were also observed (Figure 2.8).
The resistance of the connector is higher for the deeper cut (Figure 2.7.c). An increase of the resistance of about 50% was observed when the cut was deeper. The correlation between notch depth and the maximum load Fmax was almost linear. The notch depth had a more significant effect on maximum load than the effect of heel length in Figure 2.7.c. The coefficient of variations of mean data points of the maximum load was considerably smaller than the other two responses (i.e., modulus K1 and K2). It meant that the experimental measurement of stiffness was difficult, and the maximum load of the notch would be more straightforward to be predicted by the variable of notch depth. A shallower notch connector tended to have the loaded edge crushed rather than the shear-off lamellas (cf. Section 3.7). Hence, the curves of these specimens had a more prolonged post-peak displacement that ranged from 10 to 15 mm. Optimization of the notch depth will have to balance between the performance and the post-peak behavior.
For a CCC notched connector, Jiang et al. [6] reported a serviceability stiffness per 25 mm notch-depth of 15.3 kN/mm2 and the resistance per depth of 7.1 kN/mm, while the corresponding results of our study were 12.5 kN/mm2 and 7.0 kN/mm. Furthermore, the notch connectors in this study featured rounded corners at the loaded edge, while Jiang et al. tested a full-width square notch. This detail generated a transverse component of the applied force exerted on the notch. It might be the reason for the less stiff connector observed in this study.
Table of contents :
1. LITERATURE REVIEW: VIBRATION OF CROSS-LAMINATED TIMBER – CONCRETE COMPOSITE FLOORS
1.1. TIMBER-CONCRETE COMPOSITE
1.1.1. Composite action
1.1.2. Effective bending stiffness of the composite section
1.1.3. Constituent materials
1.1.4. Shear connector
1.2. VIBRATION OF TIMBER FLOORS
1.2.1. Vibration theory
1.2.2. Parameters of vibration
1.2.3. Human perception toward floor vibration
1.2.4. Design criterions
1.3. EXPERIMENTAL METHOD
1.3.1. Experimental studies on CCC structures in literature
1.3.2. Shear tests on connectors
1.3.3. Vibration tests on beams
1.4. CONCLUSION
2. PERFORMANCE OF NOTCH CONNECTOR FOR CLT-CONCRETE COMPOSITE FLOORS
2.1. INTRODUCTION
2.2. SPECIMEN GEOMETRY AND MATERIAL
2.2.1. Materials properties
2.2.2. Test specimens
2.2.3. Test setups
2.3. EXPERIMENTAL RESULTS AND DISCUSSION
2.3.1. Overview
2.3.2. Influence of heel length
2.3.3. Influence of notch depth
2.3.4. Influence of concrete thickness and screw length
2.3.5. Influence of loading sequence
2.3.6. Influence of moisture content of timber
2.3.7. Failure types
2.4. FINITE ELEMENTS MODEL VALIDATION
2.4.1. Materials
2.4.2. Models
2.4.3. Results
2.5. CONCLUSION
3. VIBRATIONAL BEHAVIOR OF CLT-CONCRETE COMPOSITE BEAMS USING NOTCHED CONNECTORS
3.1. INTRODUCTION
3.2. MATERIALS AND METHODS
3.2.1. Materials properties
3.2.2. Specimens
3.2.3. Supports
3.2.4. Test procedure
3.3. ANALYTICAL AND NUMERICAL MODELING
3.3.1. Bending stiffness of bare CLT panels
3.3.2. Bending stiffness of CCC beams
3.3.3. Finite elements model of CCC beams
3.4. RESULTS
3.4.1. Static deflection tests of CLT and CCC beams
3.4.2. Vibration tests of CCC and CLT beams
3.4.3. Bending stiffness of bare CLT panels
3.5. ANALYSIS AND DISCUSSION
3.5.1. Bending stiffness of bare CLT panels
3.5.2. Vibration characteristics of CCC beams
3.5.3. Comparison to other studies on the CCC beams
3.6. CONCLUSION
4. OPTIMIZATION MULTI-OBJECTIVE OF CLT-CONCRETE COMPOSITE FLOORS USING NSGA-II
4.1. INTRODUCTION
4.1.1. Structural multi-objective optimization
4.1.2. Multi-objective optimization algorithm
4.2. CLT-CONCRETE FLOOR DESIGN
4.2.1. CLT-concrete composite floor and the reference design
4.2.2. Notch connector influence
4.2.3. Design constraints
4.2.4. Optimization variables
4.3. OPTIMIZATION PROBLEMS
4.3.1. Objectives functions
4.3.2. Constraints functions
4.4. OPTIMIZATION ALGORITHM
4.5. ANALYSIS AND DISCUSSION
4.5.1. Pareto front of the optimal solutions and the reference solution
4.5.2. Parametric study
4.6. CONCLUSION
CONCLUSION AND PERSPECTIVES
REFERENCES
