In recent years, many reinforced concrete buildings have passed 50 years after construction, and it is necessary to determine whether to continue using them or dismantle them in the future. The durability of reinforced concrete buildings is evaluated by the depth of carbonation of concrete, which is a factor that causes rebar corrosion, from design to maintenance. However, in indoor members to which water is not supplied, there is a survey result that even if carbonation of concrete progresses, rebar corrosion does not progress. The purpose of this study is to propose a new durability evaluation method for reinforced concrete buildings considering the progress of rebar corrosion after carbonation of concrete. In this report, the carbonated concrete specimens with different concrete cover thickness of 10 to 30 mm ware exposed to several different humidity environments for 2 to 3.5 years and measured rebar corrosion rate after carbonation electrochemically. From the experimental results, the relationship between the water content of concrete and the corrosion rate of rebar was examined. In addition, the calculation results of the prediction of durability of reinforced concrete structures considering the progress of rebar corrosion after carbonation ware shown. The results obtained in this study are summarized below.
1. The water content of concrete was indicated by the degree of saturation of concrete, and the relationship between the saturation of concrete and rebar corrosion rate was discussed.
2. The rebar corrosion rate after carbonation was measured electrochemically.
3. When the concrete cover thickness is 10 mm, if the saturation of concrete is lower than 50%, it is evaluated as equivalent to the rebar corrosion rate after carbonation of the passive state. If it is higher than 50%, the rebar corrosion rate become high. When it is just 50%, the corrosion rate of rebar after carbonation become the maximum. It is considered that the content of water and oxygen in the pores of the concrete at the position of the rebar has an effect.
4. When the concrete cover thickness is 20 or 30 mm, the corrosion rate of the rebar after carbonation tends to increase as the water content of the concrete increased, but it is determined that the corrosion rate is generally low.
5. Based on the results of this experiment, the lifetime of reinforced concrete structures due to the progress of rebar corrosion after carbonation was calculated. When the saturation of concrete exceeds 50%, it is about 9 to 22 years. When the saturation of concrete is less than 50%, it is about 120 to 190 years.
6. The lifetime was calculated considering both the progress of concrete carbonation and rebar corrosion. As a result, regardless of the exposure environment, when the concrete cover thickness is 10 mm, it is over 90 years and when the concrete cover thickness is 20 mm and 30mm, it is over 200 years.
7. The reasonable durability evaluation method for reinforced concrete buildings considering both the progress of concrete carbonation and rebar corrosion in the exposed environment was presented.
Japan is located in the warm and wet climate zone. Therefore, from the standpoint of visual assessment of landscape design, it is important to take care of the color changes of building and civil engineering materials in the rain. It is well known that hygroscopic solids, such as concrete, show wet colors; Their lightness values become lower and their chroma values become higher when they get wet. A lot of results of wet color measurements were reported in the past. However, these previous reports measured only two states, dry and wet. There are no detailed reports of color changes during wetting or drying process.
Suppose that there is a material that shows wet color. It is expected that its color change during wetting or drying process will depend on the degree of its wettability. This means that, two materials showing the same wet colors in the rain may show different colors after the rain stopped, i.e., during drying. This information will be necessary and useful in the landscape design under changing weathers.
In this paper, we measured the color change during drying of beige-, red- and brown-colored garden bricks. In the experiment, we soaked the bricks in distilled water for 30 minutes to wet them completely. Then, after taking them out from water, we measured their color changes during drying within 1 minute, at five-minute interval to 60 minutes, at ten-minute interval to 90 minutes, and at 120 minutes. In the experiment, we let bricks dry naturally at room temperature 23±1 ℃ and relative humidity 50±5 %.
The results are as follows: The drastic color change arose at 5 minutes for red brick, and between 15 and 20 minutes for beige brick. For brown one, the gradual color changes occurred until 40 minutes. These color change differences during drying can be possibly explained by their differences of water content ratio.
In the wake of major earthquakes such as the 1995 Hyogo-ken Nanbu Earthquake and the 2011 off the Pacific coast of Tohoku Earthquake, it becomes very important how to design response control structures. In many cases, the response control of buildings such as winds or earthquakes mainly adding stiffness elements or adding damping elements to absorb vibration energy. On the other hand, in recent years, a device that generates an inertial resistance according to the relative acceleration between two mass points has been proposed, and a response control design using an apparent "dynamic mass (D.M.)" is possible. In the past research, there are two types of design methods using the dynamic mass (D.M.): “Mode control” and “Tuned mode control”. In this paper, we propose the M-CK type tuned dynamic mass system based on “Tuned mode control” and aim at the establishment of a more efficient response control structure.
The findings obtained from this paper are shown below.
(1) We theoretically derived M-CK type optimal design formulas based on fixed point theory and complex eigenvalue problem. Since optimal design formulas are relational formula of natural periods, it can be applied to a multi-DOF system model analysis model by using complex eigenvalue analysis.
(2) In the multi-DOF system model using the M-CK type, it was found from the results of complex eigenvalue analysis that viscous damping was imparted not only to the mode-tuned 1st mode but also to higher modes. In addition, it was shown from the amplification factors that there was an effect of reducing the response of the higher modes.
(3) The outline of the vibration test using the 8-story shear model and the results were shown. In the case of sine vibration testing, the result of the relative displacement amplification factors confirmed that the response of the higher modes was reduced by the M-CK type, and that the analytical values and the experimental values generally corresponded well. In addition, in the case of ground motion vibration testing, the effect of reducing the maximum response acceleration by the M-CK type against the non-seismic model was also experimentally verified.
(4) By using optimal design formulas and the optimal design method of the M-CK type proposed in this paper, it can be applied to full-scale buildings such as high-rise buildings. It was also shown that a more efficient control structure could be constructed.
As future developments, we will show an example of studying the optimal placement position and higher order modes control of the M-CK type. Furthermore, we show the response control effect of the combination of different tuned dynamic mass systems of the MC-K type and the M-CK type, and examine the possibility of constructing more efficient response control structure.
This paper discusses tuned mass damper (TMD) for reducing seismic response in reinforced concrete buildings. Optimal tuning ratios are estimated using equivalent two-degree-of-freedom systems through time history analysis. A formula for calculating the optimal tuning ratio is proposed in terms of the envelope curve of the restoring force and ductility factor of buildings. This formula is applicable to buildings with a wide range of natural period and soil classes for historical ground motions with an inherent complex response spectrum. Finally, a numerical example is demonstrated for buildings subjected to long-period ground motions to show the effectiveness of TMD.
In recent years, tuned mass dampers (TMD) have been applied to steel buildings for reducing seismic response in Japan. Although TMDs are also expected in reinforced concrete buildings, a tuning strategy is difficult for such highly nonlinear structures. This paper discusses the optimization of TMDs applied to nonlinear structures using time history analysis.
In Chapter 2, we propose an equivalent two-degree-of-freedom (2-DOF) system for computationally efficient optimizations. A combination of modal coordinate and physical coordinate is employed for a TMD connected to a building. To meet a compatibility condition, we present a computational method with a modified envelope curve of building hysteresis and a quasi external force applied to the TMD. This technique allows commercial software to compute the response of the 2-DOF system.
In Chapter 3, it is shown that the proposed 2-DOF system gives almost the same response with a corresponding original vibratory system. This 2-DOF system demonstrates response surfaces consisting of design parameters of a TMD, using a building having twenty-five stories. If an elastic-based optimal damping ratio of a TMD is used, control performance is approximately the same as the exact optimal TMD. This study, therefore, focuses on a tuning ratio, which specifies optimal stiffness.
In Chapter 4, optimal tuning ratios accompanied with gradually amplified ground motions are estimated using time history analysis with simulated ground motions. We formulate a closed-form of optimal tuning ratios based on equivalent natural period and a mass ratio, which is a relative mass of the TMD to the building. If the envelope curve of a controlled building and is specified, the optimal tuning ratio is obtained according to the peak ductility factor. Thus, an optimal TMD depends on seismic intensity. This formula is applicable to a wide variety of yield ratios, natural period, and the predominant period of the soil.
In Chapters 5 and 6, it is shown that the proposed formula is also applicable to buildings subjected to historical ground motions and long-period ground motions that are generated by assuming a Nankai-Trough earthquake.
Findings obtained by this study allows us to estimate an optimal TMD without iterative time history analyses. It is also useful to assess the seismic effectiveness of a TMD applied to a reinforced concrete building.
In recent years, it has become clear that higher levels of ground motion need to be considered in building design, as demonstrated by the Pacific Coast of Tohoku Earthquake. In addition, from the perspective of Business Continuity Planning, society is demanding high value-added buildings with seismic isolation and improved safety and seismic margins. In this paper, we propose a new seismic isolation frame with multiple seismic isolation layers with base-isolated core that penetrates all layers of the building. We set certain performance targets for this structure – acceleration less than 100 cm/s2 for all stories of a building and seismic isolation layer deformations less than the guaranteed performance of the isolation material – under Level 2 earthquake ground motion. We study the seismic isolation parameters that can effectively reduce the response of the new frame using a two-degree-of-freedom model. Moreover, we create a multi-mass system model and perform seismic response analysis. For the seismic isolation parameter, we use the numerical values calculated by the study using the two-degree-of-freedom model. The obtained findings are as follows;
(1) By complex eigenvalue analysis, it was possible to increase only the first-order damping factor without overdamping the second-order damping factor by increasing the attenuation of the base seismic isolation layer under the core. In addition, by examining the frequency transfer function, the proposed frame showed a significant response reduction effect in the resonance region compared with the multi-layer frame without the seismic isolation layer under the core.
(2) We formulated the resonance curves for the two-mass model of the proposed frame expressed in the relative coordinate system, the absolute coordinate system. In addition, we showed the eigenvector ratio that minimizes the maximum response magnification of lower mass point of the two-mass model by the amount of attenuation added to each seismic isolation layer.
(3) It was shown that the proposed frame can reduce the response magnification more than the case where the optimal tuning condition of TMD based on fixed point theory was applied to the double layer frame. At this time, the optimal equivalent eigenvector ratio was about γ = 2 regardless of the mass ratio, and the effective range in the proposed frame was shown as 1.5≦γ≦2.0.
(4) According to the time history response analysis using multi-mass system model, the response acceleration of the all layer of building was 100 cm/s2 or less and the interlayer deformation angle was 1/400 or less for the ground motion based on Level 2. In addition, it was confirmed that the response of the ground motion exceeding Level. 2 was reduced compared to the conventional seismic isolation frame.
We presented a novel semi-active oil damper that developed to break through the limitations of existing oil dampers by introducing a unique energy recovery system in our previous paper. This energy recovery damper is equipped with an auxiliary oil tank outside the main cylinder, and the oil flows between the cylinder and the tank are controlled by changing valves. Existing oil dampers, including variable damping types, always change the vibration energy to heat. However, the energy recovery damper recovers the vibration energy as strain energy of the oil in the tank and reuses it at an optimum timing to improve control efficiency.
This paper describes a possibility of enhancing the energy dissipation capacity of the above-mentioned energy recovery damper by introducing a multistep energy recovery system. In order to investigate the energy recovery process in detail, it found that there is an unrecoverable energy with a single oil tank. By adding extra oil tanks, this unrecoverable energy can be used to improve the damper’s performance. First, we examined using double energy recovery systems. Fig. 6 shows the basic configuration and mechanical model of the oil damper with double energy recovery systems and Fig. 8 shows the operating process. As shown in Fig. 10, double step energy recovery operation can improve the energy dissipation capacity. The effects of tank stiffness ratio and tank usage sequence are also quantitatively evaluated. Next, in order to generalize the multistep effect, the energy dissipation capacity using any n tanks is quantitatively provided. Eq.(23) is the dissipated energy per cycle in steady state. Based on Eq.(23), the relationships between the energy dissipation capacity and the tank stiffness ratio in steady state are shown in Fig. 11 and Fig. 12. From these figures, the effect of a total volume of oil tanks and number of energy recovery system can be observed. The energy dissipation capacity using multistep control for harmonic excitations or earthquakes are examined through a theoretical approach and numerical response analysis.
Following results are obtained by the theoretical and numerical studies.
1. There is unrecoverable energy in the operation process of the energy recovery damper. The most important point of improvement is how much unrecoverable energy can be reduced.
2. Usage of the additional tanks and control valves can reduce the unrecoverable energy and the energy dissipation capacity can be improved.
3. The theoretical study based on the mechanical models showed quantitatively the relationship between the total volume of the oil tanks and the multistep energy recovery system. Regardless of the total tank stiffness ratio that is defined by the total volume of the oil tanks, the greater the number of the oil tanks, the greater the energy dissipation capacity.
4. In the multistep energy recovery process, the order of the oil tanks to be used is better last-in first-out than first-in first-out and uniform tank stiffness ratios are better than non-uniform tank stiffness ratios.
5. The enhancing effect of the energy dissipation capacity by the multistep control confirm for harmonic excitations and earthquakes.
Soil-structure interaction (SSI) influences the seismic response of buildings supported by pile foundation. Under a significant input motion that causes damage to piles, the SSI is influenced by the nonlinear behaviors of the free ground (i.e. site nonlinearity), soil close to the foundation (i.e. local nonlinearity), and structural members (i.e. structure nonlinearity). The interaction among piles (i.e. the effect of pile groups) complicates local nonlinearity when two or more piles are installed in a narrow space. Experimental and analytical studies have indicated that lateral resistance per pile in a pile group tends to be smaller than that of a single pile, and depends on its location. However, few studies have conducted on lateral resistance of pile groups under load in the oblique direction. Therefore, this study aims to investigate seismic response of pile groups subjected to input motion in the oblique direction.
The study consists of two steps. In the first step, a shaking table test and its simulation analysis are conducted to investigate influences of direction of input motion on dynamic behaviors of a structure supported by a pile group. In the second step, a cyclic loading analysis on full-scale pile groups is conducted to evaluate hysteresis characteristics of pile-soil springs. The analyses are based on the nonlinear three-dimensional finite element method.
The shaking table test was for a rigid body supported by a 5 x 5-pile group. Piles were modeled using acrylic cylinders with a diameter of 12 mm and a length of 421 mm, and installed in the dry Toyoura sand deposit. Six types of input motion, which contains two different waveforms and three different amplitudes, were applied to 0-degree or 45-degree direction. The experiment provided following findings:
1) The amplification characteristic of the soil-structure system did not depend on direction of input motions. This fact indicates lateral resistance of a whole of the pile group does not depend on direction.
2) Subgrade reaction of each pile differed significantly depending on the pile location and the direction of input motion. This difference was considered to cause that of bending moment.
The cyclic loading analysis was for a 2-pile group, 3-, and 5-pile groups in the series arrangement, and 3 x 3- and 5 x 5-pile groups in the square arrangement. Piles had a diameter of 600 mm and a length of 10 m. Soil was assumed to be homogeneous sand deposit, and modeled using perfect elasto-plastic bodies governed by the Mohr-Coulomb yield criterion. The amplitude of pile displacement was gradually increased, and the loading direction was varied at the interval of 15 degrees. The analysis provided following findings:
1) Maximum subgrade reaction and hysteresis dissipated energy of each pile strongly depended on the loading direction. The relationship between these characteristics and loading direction differed significantly depending on pile location.
2) Equivalent damping ratio of the equivalent single pile did not depend on the loading direction as long as nonlinearity in soil close to piles progressed.
Simulated earthquake motions whose response spectra correspond to the multi-target response spectra are often used to design earthquake-resistant nuclear power plant facilities. Therefore, it is advantageous to utilize simulated ground motions with an envelope curve of actual earthquake ground motions while designing these facilities. These motions can be generated by several methods. However, these methods cannot maintain the envelope curve of the simulated earthquake motions. A method to generate simulated earthquake motions that satisfies both the multi-target response spectrum and the envelope requirements does not exist.
This paper proposed a novel method for generation of simulated earthquake motions by considering an actual earthquake envelope curve and multi target response spectra. This method is equivalent to the modal iterative error-correction method that is effective for solving inverse problems with significant discontinuities. Therefore, this paper proposed an improvement of this method for envelope curve and spectrum fitting.
The findings of this study are as follows:
(1) A novel method for generation of simulated earthquake motions that considers actual earthquake envelope curve and multi-target response spectra was developed. In this method, the input vector was changed from the amplitude spectrum to the acceleration vector of input motion. In addition, the targets were both the spectra and the envelope curve. Subsequently, the acceleration vector of input motion was changed (while maintaining a constant envelope curve) by using the modal iterative error-correction method to fit the target spectra and target envelope curve.
(2) The proposed method was applied to sample problems to verify its accuracy. The results indicated that the simulated seismic motion accurately represented the envelope curve of the actual earthquake. Additionally, two acceleration response spectra were also created.
(3) However, depending on the setting of the envelope curve, there was no suitable simulated ground motion. In such a case, it is effective to make a simulated earthquake motion by relaxing the restriction of the envelope curve, which is realized by increasing the time width of the envelope curve.
This paper presents a numerical method for three-dimensional (3D) seismic response analysis of a frame containing member failure. The base of proposed method is the Fibered Plastic Hinge Model (FPHM)9),10) in which elastic and plastic components of deformations of each element can be separated explicitly from a largely deformed frame. The analysis is performed with cancellation of a large unbalanced force vector caused by a sudden fracture of members in the dynamic elastoplastic incremental analysis of a frame. The FPHM program uses a gradient of existing elastic strain energy as an internal force vector, which is needed to evaluate unbalanced force vector, of a frame at each incremental step. A redistribution of member forces, which are axial force, biaxial bending moments, shear force and axial torsional moment, of early fractured low ductile member into the remaining members of the frame is done in each step, therefore, the dynamic response of a frame that contains both low ductile and ductile members can be obtained accurately. The validity of proposed method is verified through the numerical experiments on one-bay one-story braced steel frame having a low ductile tension brace.
Then, a possibility to use proposed method as a collapse analysis method for a 3D frame is examined by utilizing available shaking table test results on full-scale two-bay four-story steel building13) assuming a simple fracture criterion for an element:
|ε|max = ηεy
where |ε|max is the maximum value of axial strain of a fiber due to varying axial force and biaxial bending moments at the element ends, εy is the initial yield strain of a fiber, and η is a reference value. Since the FPHM divides the element-end sections to fine fibers, |ε|max can be easily obtained in the numerical procedure.
Assuming η = 20, which was determined by trial and error, the obtained numerical results follow mostly the collapse behavior of the building, except that the deterioration behavior due to local buckling of the columns is different from the test result13). In addition, the firstly and secondly fractured columns obtained by the present analysis are consistent with those observed in the test13). Although a systematic way to estimate the value of η is unknown at the present time, η may be a parameter which relates a member which loses load carrying capacity by the local buckling to an element formulated according to the Bernoulli-Euler hypothesis.
Question of whether it is necessary for stress method to use the condition for compatibility of strain or not might be brought up from the fact that the stress can derive from a formulation based on stress method by means of application of the property of the Moore-Penrose generalized inverse without the help of the compatibility condition.
One answer to the question might be found in an analysis of the situation where the reciprocal theorem is not satisfied because of the reason that the transposed relation between the stress-force matrix and the displacement-strain matrix is disregarded. The situation might be able to find in the FEM analysis as rare case where the stress-nodal force matrix of element is independently made up without considering the corresponding relation of the nodal displacement- strain matrix.
When the reciprocal theorem is not satisfied, we will face two kinds of problem to solve. One is to clear up the definition and the role of what is called the condition for compatibility of strain which is usually used to decide redundant force. Another is how to derive displacement from stress without the help of the principle of energy like the Castigliano’s theorem. Answer of each problem and the explicit forms of stress solution and displacement solution are represented in this paper.
Results can summarize as follows,
(i) What is called the condition for compatibility of strain, which is re-defined as the condition for existence of strain to the compatibility equation relating displacement to strain in paper, is equivalent to the orthogonal condition between strain and self-stress when the reciprocal condition is guaranteed. Which means it is necessary to clear up the distinction between the existence condition and the orthogonal condition when the reciprocal condition is not satisfied. The distinction is available to obtain stress as shown in the section 5.1.
(ii) Displacement is obtained by the inverse operation of the generalized inverse matrix as the particular solution of the compatibility equation as shown in the section 5.2.
(iii) Merit of the explicit forms of stress solution and displacement solution obtained in this paper is easily provable to coincide with the usual solution forms of displacement method. Origin of the merit is easy operation of the orthogonal projection matrix expressed by the generalized inverse. Then stress variables can treat as a set instead of a choice of the independent component of stress variables such as redundant force needed to make up the regular matrix.
(iv) There is nothing to change the formulation of getting displacement and stress in case of displacement method even if the reciprocal theorem is satisfied or not. On the other hand, when the reciprocal theorem is kept in stress method the computational complexity of getting the solutions drastically decreases in less level than the corresponding displacement method Therefore, the reciprocal theorem should be kept to the stress method. How to apply the solution forms to the FEM analysis based on stress method is illustratively shown in example 2 where a square structure of the state of plane stress is divided into two triangular elements. The same structure composed of single rectangular element where the stiffness matrix is not regular is also treated as shown in example 3, and so on.
Buckling is a critical problem for column material. The Japanese Building Standards Law prescribes the upper limit of the slenderness ratio for wood materials at 150. This upper limit is based on the result of lumber evaluations; therefore, it might not be suitable for engineered wood, such as glulam or LVL. Recently, large or medium-sized wooden buildings have attracted much attention, and larger wooden buildings have been built in various regions of the world. The buildings with long columns whose slenderness ratio is over 150 are attractive in Japanese building market.
In previous studies, one of the authors suggested a method to evaluate the buckling strength based on experimental data using LVL for brace in structural walls and plywood. The cross-sectional shape was rectangular and limited.
In this study, we evaluated the buckling strength using glulam and LVL for columns. The cross-section was square, i.e., 120 × 120 mm and 60 × 60 mm. The slenderness ratio was set between 30 and 200, and the laminated direction for deflection was also set as a parameter. These specimens were evaluated using compression tests, where both ends were pinned-support. The bending Young's modulus was measured by a non-distractive vibration test before conducting the compression tests. We conducted monotonic loading tests with short columns to define the yield strain for the evaluation of buckling strength. Regarding the short columns, the yield strain of each material was defined as 1300–1400 μ for the glulam specimen and 1400–1500 μ for the LVL. The limitation slenderness ratio was calculated using these values. From the vibration tests, Young’s modulus of glulam was 10.31–10.39 kN/mm2, and that of LVL was 13.75–13.82 kN/mm2.
Three specific values were used to evaluate the buckling strength of the specimens, namely maximum stress (σ1), stress level evaluated by the Southwell method (σ2), and plastic buckling strength (σ3). The values of σ1 and σ2 were similar; however, the conclusion was that σ1 was more appropriate for the evaluation of safety. By comparing σ3 and Euler's buckling strength formula, and comparing σ1 and σ3, we determined that the critical slenderness ratio of glulam and LVL was 100. This value corresponds with Standard for Structural Design of Timber Structures, and the basis of the standard value can be shown more theoretically. Furthermore, the comparison of the Euler buckling strength formula and test results proved that the buckling strength can be evaluated safely for materials with a slenderness ratio of up to 200. Regarding the suitability of Young's modulus for Euler’s formula, we found that we could use a 5% lower limit value of the material tests or those in JAS for glulam and LVL. The test results and calculated values correlated well.
Recently, from view－point of Global Environment, timber, i.e., one of nature-cycle materials, has been tried to be utilized as structural members of large timber buildings in Europe and North America. A representative timber of the members is Cross-laminated timber (CLT), however, CLT structural system very often restricts planning of building because of CLT being plate member. High-stiffness-strength-timber slender beam and column are significantly desired.
This study focuses a hybrid glulam timber beam with steel deformed bar (rebar) and Epoxy resin adhesive. Aim of this study is to derive equations for internal stress of hybrid glulam timber beam by temperature and moisture content variation in its timber were derived from a differential equation. The result of derivation, numerical analysis, and its discussion are summarized as follows:
(i) A spring model for the hybrid glulam timber beam was proposed to derive the equations, by simplifying essential elements of the hybrid beam.
(ii) Effects on internal stress when temperature rises 60℃ or moisture content(MC) rises 3% were discussed; equations for axial stress of rebar and timber in way of beam axis, and bond stress around rebar at MC of 3% were suggested.
(iii) The equations were verified with numerical analysis for two beam cross sections with span lengths from 4-m to 20-m. Calculations by the equations were fairly good agreed with calculations by the numerical analysis.
(iv) Under natural environment in Japan, air temperature rises with air humidity, therefore effects on internal stress in the hybrid beam of timber temperature and moisture content are estimated to decrease. Also, large size timber utilized for the hybrid beam could not absorb air moisture compared with smaller size timber. Hereafter, standardization of temperature and moisture content for the hybrid beam will be required.
When a new member is connected to an existing concrete member, roughened concrete and post-installed anchors are often applied to the joint. An electric hammer is used to create the uneven roughened concrete surface. There appears to be no detailed and general design code with regard to the shape and strength of the concrete in existing design guidelines. However, according to a series of studies conducted by us, setting the area ratio to 0.3 or more is important for shear resistance in practical constructions. However, there is a problem of how to manage the uneven shape that differs considerably in the consciousness and technique of contractors. For this purpose, a tool that can easily calculate the roughened area ratio quantitatively at construction sites is required. In recent years, deep learning has been applied to various studies and applications. Deep learning does not require new advanced algorithms to be designed and can make complex decisions quantitatively. Therefore, in this study, we investigated whether the roughened area ratio could be measured from photographs using deep learning. Thus, we investigated whether the roughened area can be classified using deep learning and what learning conditions ware suitable.
First, specimens with roughened surfaces were prepared and trained via deep learning. Hence, it is possible to classify the roughened surface of the test specimen in a satisfactory manner; however, classifying the roughened surface of the practical concrete members is not possible. Because the roughness of the specimen is different from that of the practical frame, it is assumed that training with a practical frame is required to apply it to a practical frame.
Subsequently, the images of the practical frame were trained. Here, photographs of the two buildings were prepared, and the combination was changed to verify the performance of deep learning based on the difference in the tendency of the training data. As a result, training that involved images with various roughened shapes and smooth surface conditions could cope with various types of roughening. On the contrary, in training using images with uniform roughened shapes and uniform smooth surface conditions, the classification performance toward other types of images significantly reduces, although the classification performance toward similar images is high. For this reason, the performance of deep learning largely depends on the training data. Therefore, performing training using data suitable for the required performance is preferable.
Subsequently, we investigated the conditions of deep learning to increase the precision such that the measured roughened area ratio was as safe as possible. As a result, the highest precision was obtained by applying DeepLab v3+, Inception-ResNet-v2, and SGDM for stochastic gradient descent. On the contrary, under some bad conditions, the roughened surface could not be evaluated. As mentioned above, evaluating the roughened surface by appropriately training roughening using deep learning is possible, and we believe that it can contribute to site management.
Transverse stiffeners are effective for lateral buckling of H-shaped steel beams with continuous restraint on the upper flange because of constraining the deformation of the cross-section of beam members. There are, however, few studies on the additional stiffening effect of stiffeners on continuously restrained beam members, and the position to install stiffener is not clear. Therefore, the purpose of this paper is to evaluate the buckling strength of H-shaped beam members with continuous restraint and installed transverse stiffeners — furthermore, the position where stiffeners should be installed is investigated.
In this study, the elastic buckling strength is calculated by the theoretical analysis using the energy method, and from these numerical simulations, the relationship between the geometric shape of the beam members and the stiffening effect is clarified. Also, the evaluating equation for lateral buckling strength is established by using the numerical simulation results. Moreover, the numerical analyses based on the finite element method are performed. The classification of the collapse mode, the maximum strength and the plastic deformation capacity are examined.
From this research, the following are found.
1) The stiffening effect (i.e., the increase of lateral buckling strength) is more effective at a position where the deformation of lateral buckling of the beam members without stiffeners is significant. In the case of stiffeners are installed at only one place on beam members, the stiffening effect increases as the moment gradient becomes large and the compressive region of the lower flange becomes narrow. Also, the cross-sections which have a large effect on the restraining upper flange are hard to deform due to lateral buckling, so the stiffening effect is also hard to increase.
2) New evaluating index λcr was proposed. By using new index λcr, it is possible to evaluate the stiffening effects regardless of the cross-sectional shape. By means of the new index, moreover, the equation (3.4), which evaluates the elastic buckling strength of the H-shaped beam members with continuous restraint installed transverse stiffeners, has been proposed.
3) Optimum stiffening position LS* is defined as the position where maximizes the stiffening effects. The equations evaluating optimum stiffening position and the maximum stiffening effect are derived by approximating.
4) The elastic buckling strength of H-shaped beam members with continuous restraint and installed transverse stiffeners can be calculated based on equation (3.4). When the coupled buckling occurs, however, a slight reduction in the buckling strength should be considered. Especially in the case of the stiffening position is near the end of the beam members, the buckling strength is significantly reduced.
5) The classification of the collapse mode is possible by means of the generalized slenderness ratio λb, which is modified by the evaluating formula (3.4) for the lateral buckling strength. Furthermore, the evaluations of the maximum strength and the plastic deformation capacity are also possible by means of the generalized slenderness ratio λb.
The objective of this study is to present the expressions for calculating the ultimate flexural strength and to examine the relations between the superposed strength and the ultimate strength of rectangular concrete filled tubular sections. In addition to the superposed strength, a method to calculate the ultimate flexural strength is presented. Numerical calculation based on the theoretical analysis is performed taking the width-to-thickness ratio of steel tube, strengths of steel and concrete and the ultimate strain as the analytical parameters, and moment -axial load interaction relationships are shown. Comparing the ultimate strength with superposed strength, the effects of analytical parameters on the strengths are discussed.
2. Analytical work
The expressions to calculate the generalized superposed strength of the rectangular CFT sections shown in Fig. 1 are given as the Eqs. (8) ~ (12), together with the expressions for the simple superposed strength of Eqs. (14)~(18). In addition to the superposed strength, the expressions for calculating the ultimate strengths are shown as the Eqs. (28) ~ (32) and (34) ~ (37).
3. Results and discussions
As the analytical parameters, the width-to-thickness ratio, yield stress of steel tube, compressive strength of concrete and ultimate strain of concrete are selected, and they vary as follows; 1) width-to-thickness ratio 20 and 40, 2) yield stress of steel tube 325, 440 and 700N/mm2, 3) compressive strength of concrete 30, 60 and 90 N/mm2 and 4) ultimate strain of concrete 0.004 and 0.008.
Relationship between the non-dimensional moment and axial load are shown in Figs. 11 and 12. In these figures, group "CFT", "S" and "C" denote the strength of CFT sections, steel tubular sections and concrete sections, respectively. It is shown that the effect of the ultimate strain on CFT strength becomes large as the strengths of steel and concrete become large. Figure 13 show the relationships the ratio of simple superposed strength S+CMpc to generalized superposed strength cftMpc and the axial load ratio. It is observed that the ratio S+CMpc/cftMpc becomes large as the effect on strength by the concrete becomes small. Figures 14 and 15 show the Mu/cftMpc ratio, where Mu denotes ultimate flexural strength. In the case of ultimate strain is equal to 0.004, the minimum ratio Mu/cftMpc is about 0.7, whereas the ratio is about 0.9 when the strain is 0.008.
The conclusions derived from this study are as follows:
1) The expressions for calculating the ultimate strengths are presented as the Eqs. (28) ~ (32) and (34) ~ (37).
2) The ratio S+CMpc/cftMpc becomes large as the effect on the strength by the concrete becomes small.
3) In the case of the ultimate strain is equal to 0.004, the minimum ratio Mu/cftMpc is about 0.7, whereas the ratio is about 0.9 when the strain is 0.008.