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Steel Construction Published the List of Highly Influential Articles of 2020
2020 Vol. 35, No. 4
Display Method:
2020, 35(4): 1-10.
doi: 10.13206/j.gjgS20051202
Abstract
PDF (4168KB)
Abstract:
Transmission tower is an important load-bearing facility of transmission line, and its safety is directly related to the normal operation of the national grid and transmission line. Wind-induced response of transmission towers is mainly studied by field measurement, wind tunnel test and numerical simulation. With the development of computer technology and numerical methods, numerical simulation analysis on wind-induced response of transmission towers begins to be widely adopted and significant achievements were gained. Wind load model and structure model are established, then the structure dynamic response characteristics and the corresponding wind vibration control method are studied in related numerical simulation research, so progress of wind-induced response numerical simulation research of transmission tower is summarized from wind load model, structure model and dynamic response characteristics and wind vibration control research in this article.
The mean wind and fluctuating wind model of wind field in the ground layer is the basis of building structure wind load. The wind speed profile model used for the mean wind mainly includes exponential and logarithmic wind speed profile model, while the fluctuating wind is mainly simulated according to turbulent wind power spectrum. Under different extreme weather conditions, wind field shows different characteristics from normal wind. The corresponding mean and fluctuating wind models need to be further studied according to the actual situation. The wind load of transmission tower also needs relevant structural parameters, in which the wind resistance effect of tower structure and the shielding effect between tower components can be studied by flow field simulation.
When building the transmission tower finite element model, the transmission tower can be regarded as the rigid frame structure and the truss-beam structure, while the error of simulation by using the truss model is large. In addition to wind load and other external environmental loads, the influence of transmission line on tower structure should also be considered, so the tower-line coupling system should be established to simulate the actual structure characteristics of transmission tower. In the process of building the finite element model of tower-line system, the catenary theory and the horizontal tension of conductor can be used to model and shape the conductor.
Based on the wind load model and the structural model, the wind-induced response of transmission tower can be analyzed. The dynamic characteristics of the structure have important effects on the wind-induced response, and the effect of the conductor on the tower makes dynamic characteristics of tower-line system more complex. For the wind load of tower under different wind direction, the relevant codes have corresponding calculation coefficient and distribution coefficient. For the tower-line coupling system, the wind direction has more significant effects on the wind-induced response.
According to whether external energy input is needed, wind-induced vibration control can be divided into active control, passive control and hybrid control. So far passive control, especially tuned mass damper, is still the main method for wind-induced vibration control of transmission tower. The natural frequency of damper should be consistent with the natural frequency of tower, then the wind-induced vibration control works best. However, the optimization of wind-induced vibration control is more complicated due to tower-line coupling effect.
Besides, future research direction was prospected. Further research on wind field characteristics of special weather, development of more reliable finite element modeling methods, further study of tower torsional and along-line response characteristics, and optimization of TMD design parameters and layout should be important research directions in the future.
Transmission tower is an important load-bearing facility of transmission line, and its safety is directly related to the normal operation of the national grid and transmission line. Wind-induced response of transmission towers is mainly studied by field measurement, wind tunnel test and numerical simulation. With the development of computer technology and numerical methods, numerical simulation analysis on wind-induced response of transmission towers begins to be widely adopted and significant achievements were gained. Wind load model and structure model are established, then the structure dynamic response characteristics and the corresponding wind vibration control method are studied in related numerical simulation research, so progress of wind-induced response numerical simulation research of transmission tower is summarized from wind load model, structure model and dynamic response characteristics and wind vibration control research in this article.
The mean wind and fluctuating wind model of wind field in the ground layer is the basis of building structure wind load. The wind speed profile model used for the mean wind mainly includes exponential and logarithmic wind speed profile model, while the fluctuating wind is mainly simulated according to turbulent wind power spectrum. Under different extreme weather conditions, wind field shows different characteristics from normal wind. The corresponding mean and fluctuating wind models need to be further studied according to the actual situation. The wind load of transmission tower also needs relevant structural parameters, in which the wind resistance effect of tower structure and the shielding effect between tower components can be studied by flow field simulation.
When building the transmission tower finite element model, the transmission tower can be regarded as the rigid frame structure and the truss-beam structure, while the error of simulation by using the truss model is large. In addition to wind load and other external environmental loads, the influence of transmission line on tower structure should also be considered, so the tower-line coupling system should be established to simulate the actual structure characteristics of transmission tower. In the process of building the finite element model of tower-line system, the catenary theory and the horizontal tension of conductor can be used to model and shape the conductor.
Based on the wind load model and the structural model, the wind-induced response of transmission tower can be analyzed. The dynamic characteristics of the structure have important effects on the wind-induced response, and the effect of the conductor on the tower makes dynamic characteristics of tower-line system more complex. For the wind load of tower under different wind direction, the relevant codes have corresponding calculation coefficient and distribution coefficient. For the tower-line coupling system, the wind direction has more significant effects on the wind-induced response.
According to whether external energy input is needed, wind-induced vibration control can be divided into active control, passive control and hybrid control. So far passive control, especially tuned mass damper, is still the main method for wind-induced vibration control of transmission tower. The natural frequency of damper should be consistent with the natural frequency of tower, then the wind-induced vibration control works best. However, the optimization of wind-induced vibration control is more complicated due to tower-line coupling effect.
Besides, future research direction was prospected. Further research on wind field characteristics of special weather, development of more reliable finite element modeling methods, further study of tower torsional and along-line response characteristics, and optimization of TMD design parameters and layout should be important research directions in the future.
2020, 35(4): 19-27.
doi: 10.13206/j.gjgS19102901
Abstract:
Steel-concrete composite beams with corrugated webs have higher bearing capacity than traditional composite beams.In this paper, ABAQUS finite element method is used to established the corrugated web under the high temperature thermal-mechanical coupling model.The influence of load ratio, span-height ratio, web wave angle, corner constraint ratio and axial constraint ratio on the fire resistance of composite beams is studied by analyzing the mid-span deflection and the variation curve of axial force at beam end with temperature of corrugated web composite beams.The results show that neutral axis of composite beams with corrugated webs moves up continuously along section height and finally forms the plastic hinge and reaches the critical state when exposed to fire. Without rotation and axial restraints, the load ratio has a greater impact on the fire resistance of steel-concrete composite beams with corrugated webs; as the load ratio increases, the critical temperature of the composite beam also decreases; under high temperature, the span-to-height ratio increases. The smaller, the deflection value of the composite beam in the critical state is smaller; under other conditions being the same, the larger the web wave angle is, the closer it is to a flat web composite beam, the more likely it is that local instability will occur, and the fire resistance will be worse. After considering the beam rotation and axial restraints, the critical temperature of the composite beam is higher under the same load ratio, which is about 1.2 to 1.5 times the critical temperature of the unconstrained composite beam. When the critical state is reached, the deflection is smaller, about 0.6 to 0.8. When the corner constraint is large, The larger the corner restraint ratio of corrugated web composite beams, the slower the rate of mid-span deflection decrease, the greater the negative bending moment at the beam end, and the more prone the lower flange of the beam end to buckling failure under compression. When the wave angle of the web is 65, the mid-span deflection of the composite beam with non-end restraint corrugated web decreases fastest, while the composite beam with end restraint corrugated web has the largest deflection when the wave angle of the web is 45.
Steel-concrete composite beams with corrugated webs have higher bearing capacity than traditional composite beams.In this paper, ABAQUS finite element method is used to established the corrugated web under the high temperature thermal-mechanical coupling model.The influence of load ratio, span-height ratio, web wave angle, corner constraint ratio and axial constraint ratio on the fire resistance of composite beams is studied by analyzing the mid-span deflection and the variation curve of axial force at beam end with temperature of corrugated web composite beams.The results show that neutral axis of composite beams with corrugated webs moves up continuously along section height and finally forms the plastic hinge and reaches the critical state when exposed to fire. Without rotation and axial restraints, the load ratio has a greater impact on the fire resistance of steel-concrete composite beams with corrugated webs; as the load ratio increases, the critical temperature of the composite beam also decreases; under high temperature, the span-to-height ratio increases. The smaller, the deflection value of the composite beam in the critical state is smaller; under other conditions being the same, the larger the web wave angle is, the closer it is to a flat web composite beam, the more likely it is that local instability will occur, and the fire resistance will be worse. After considering the beam rotation and axial restraints, the critical temperature of the composite beam is higher under the same load ratio, which is about 1.2 to 1.5 times the critical temperature of the unconstrained composite beam. When the critical state is reached, the deflection is smaller, about 0.6 to 0.8. When the corner constraint is large, The larger the corner restraint ratio of corrugated web composite beams, the slower the rate of mid-span deflection decrease, the greater the negative bending moment at the beam end, and the more prone the lower flange of the beam end to buckling failure under compression. When the wave angle of the web is 65, the mid-span deflection of the composite beam with non-end restraint corrugated web decreases fastest, while the composite beam with end restraint corrugated web has the largest deflection when the wave angle of the web is 45.
2020, 35(4): 28-38.
doi: 10.13206/j.gjgS20010803
Abstract:
The reciprocal structure has a long history, which has attracted the attention of the academic and engineering circles due to its advantages of simple node, convenient construction and beautiful shape. However, the existence of a large number of geometric constraints in the reciprocal structure makes its configuration difficult, which becomes a big obstacle to the application of the reciprocal structure. In order to find feasible schemes for generating reciprocal configurations on a given surface, the feasibility of applying the geometries of Archimedean pavings to generate reciprocal configurations on the cylindrical surface is investigated. For the feasibility judgment of cylindrical grid generation of reciprocal configurations, the judgment process in this paper is to carry out the plane pre-judgment first, then directly transform the cylindrical grid, and then carry on the judgment through the three-level judgment method or the fast judgment method. 11 kinds of Archimedes plane pavements are curved with different axis of symmetry as the longitudinal axis, and 21 kinds of Archimedes cylindrical meshes can be generated. Therefore, there are 63 possible schemes for the generation of reciprocal configurations by three direct conversion methods:contraction method, element-rotating method and extended translation method.
These 63 schemes were systematically determined and screened in this paper. Firstly, the typical features of some infeasible schemes can be obtained through the plane pre-judgment and the physical judgment of nodal elements in the three-level judgment method. According to these characteristics, the grid form of Archimedes' paved cylinder can be preprocessed in batches to eliminate some infeasible schemes. For the rest of the schemes, through the developed MATLAB program based on geometric analytical solutions, transformations from Archimedean pavings to reciprocal configurations were conducted employing contraction method, element-rotating method and extended translation method, and furthermore judgment methods at three level and a fast judgment method were employed for determinations one by one. The results show that among the 63 possible schemes, there are only 6 schemes theoretically satisfy the requirements of reciprocal configuration, which are A1② by contraction method and element-rotating method, A2② by extended translation method, A3② by contraction method and element-rotating method, and A3① by contraction method. The analytical solutions of these 6 feasible schemes are given in this paper. The grids of these 6 configurations have the characteristics of single cell geometry, symmetrical rotation of the second degree and regular overall morphology. In order to validate the results of the study, practical entity reciprocal configuration models were successfully built for 5 feasible schemes.
The reciprocal structure has a long history, which has attracted the attention of the academic and engineering circles due to its advantages of simple node, convenient construction and beautiful shape. However, the existence of a large number of geometric constraints in the reciprocal structure makes its configuration difficult, which becomes a big obstacle to the application of the reciprocal structure. In order to find feasible schemes for generating reciprocal configurations on a given surface, the feasibility of applying the geometries of Archimedean pavings to generate reciprocal configurations on the cylindrical surface is investigated. For the feasibility judgment of cylindrical grid generation of reciprocal configurations, the judgment process in this paper is to carry out the plane pre-judgment first, then directly transform the cylindrical grid, and then carry on the judgment through the three-level judgment method or the fast judgment method. 11 kinds of Archimedes plane pavements are curved with different axis of symmetry as the longitudinal axis, and 21 kinds of Archimedes cylindrical meshes can be generated. Therefore, there are 63 possible schemes for the generation of reciprocal configurations by three direct conversion methods:contraction method, element-rotating method and extended translation method.
These 63 schemes were systematically determined and screened in this paper. Firstly, the typical features of some infeasible schemes can be obtained through the plane pre-judgment and the physical judgment of nodal elements in the three-level judgment method. According to these characteristics, the grid form of Archimedes' paved cylinder can be preprocessed in batches to eliminate some infeasible schemes. For the rest of the schemes, through the developed MATLAB program based on geometric analytical solutions, transformations from Archimedean pavings to reciprocal configurations were conducted employing contraction method, element-rotating method and extended translation method, and furthermore judgment methods at three level and a fast judgment method were employed for determinations one by one. The results show that among the 63 possible schemes, there are only 6 schemes theoretically satisfy the requirements of reciprocal configuration, which are A1② by contraction method and element-rotating method, A2② by extended translation method, A3② by contraction method and element-rotating method, and A3① by contraction method. The analytical solutions of these 6 feasible schemes are given in this paper. The grids of these 6 configurations have the characteristics of single cell geometry, symmetrical rotation of the second degree and regular overall morphology. In order to validate the results of the study, practical entity reciprocal configuration models were successfully built for 5 feasible schemes.
2020, 35(4): 39-49.
doi: 10.13206/j.gjgS20041302
Abstract:
Design methods of centrally loaded member for AISC 360-16 are explained and comparison is make between AISC 360-16 and GB 50017-2017. Design considerations and differences of flexure buckling, torsion buckling and flexure-torsion buckling for two country's codes are introduced. These will be helpful to understand design accomplishments for three kinds of buckling.
The stability strength calculation of axial compression member is listed in Chapter E of AISC 360-16, in which the design compressive strength is φcPn; the compressive coefficient of drag is φc=0.9; the nominal compressive strength Pn shall be the lowest value obtained based on the applicable limit states of flexural buckling, torsional buckling, and flexural-torsional buckling.
For flexural buckling, the number of column curves in AISC 360-16 decreases from three to one. The flexural buckling strength of axial compression member is Pn=FcrAg, where there are Fcr=0.658Fy/FeFy or Fcr=0.877Fe, and Fe=((π2E)/((Lc/r)2)).
The column curve in AISC 360-16 can be divided into two segments, with the limit slenderness ratio of 137 (Fy=235 MPa) and the normalized slenderness ratio of λc=(Lc/πr)√Fy/E=1.5. In the elastic stage, the Euler's critical force is taken as the reference, and the reduction factor 0.877 is introduced to consider the influence of the geometric imperfections of the member, which is clear in concept and consistent with Euler's formula. In the inelastic stage, the nonlinear properties of the generalized Euler's formula are taken into account after the material strength exceeds the proportional limit, and the geometric imperfections and the residual stresses of the member are also considered.
GB 50017-2017 adopts four column curves, flexural buckling strength of which can be calculated by N/φAf ≤ 1.0. Considering the initial imperfections and the residual stresses of the member, stability factor φ is determined by the critical force of flexural buckling of axial compression member and is applicable in both elastic and inelastic zones. The slenderness ratios in both main axes are first calculated and then the larger slenderness ratio is substituted into φ=(1/2λn2)[(α2+α3λn+λn2)-√(α2+α3λn+λn2)2-4λn2] or i>φ=1-α1λn2 to calculate stability factor φ.
When analyzing torsional buckling and flexural-torsional buckling, torsional buckling may occur for biaxial symmetric cross-section. For cross-sections with single symmetric axis or without symmetric axis, flexural-torsional buckling may happen. The buckling strength of torsional buckling or flexural-torsional buckling is Pn=FcrAg, where there are Fcr=0.658Fy/FeFy or Fcr=0.877Fe, and Fe takes the critical stress of torsional buckling or flexural-torsional buckling.
Thus it can be seen, torsional buckling and flexural-torsional buckling of AISC 360-16 are calculated according to the column curve of flexural buckling after the computation of the corresponding elastic critical force.
The critical force of torsional buckling in GB 50017-2017 is Nz=(1/i02)(GIt+((π2EIω)/l2)), by which divided A is consistent with that in AISC 360-16.
In specific applications, GB 50017-2017 adopts the concept of equivalent slenderness ratio, and the slenderness ratio for torsional buckling is calculated by λz=√I0/(It/25.7+Iω/lω2) and considered as the slenderness ratio for flexural buckling. The stability factor φ is then computed based on the above method, which is the same as the method adopted in AISC 360-16. The difference is that the elastic segment of column curve in AISC 360-16 is determined by the product of Euler's critical load and reduction factor, and the inelastic segment is closer to the test curve, while GB 50017-2017 takes the regression curve from test data as the column curve.
GB 50017-2017 adopts (NEy-Nyz)(Nz-Nyz)=Nyz2(ys2/i02) to calculate the critical flexural-torsional buckling force Nyz. In AISC 360-16, supposing xo=0, flexural-torsional buckling equation of uniaxial symmetric cross-section is (Fe-Fey)(Fe-Fez)=Fe2(yo/[ro)2. Essentially, the above two equations are consistent. In other words, the calculation formulas of flexural-torsional buckling strength in GB 50017-2017 and AISC 360-16 are the same.
For axial compression member, both AISC 360-16 and GB 50017-2017 present design formulas for flexural buckling, torsional buckling, and flexural-torsional buckling. Both the two specifications determine the stability factor φ of three buckling modes based on column curve of flexural buckling. AISC 360-16 adopts one column curve and GB 50017-2017 uses four column curves, both of which consider the influences of geometric imperfection and residual stress and agree well with relevant tests in different countries. Torsional buckling and flexural-torsional buckling in AISC 360-16 and GB 50017-2017 come from the elastic stability equilibrium equation, which are transformed into the critical force of flexural buckling through equivalent critical force and designed by column curve of flexural buckling.
Design methods of centrally loaded member for AISC 360-16 are explained and comparison is make between AISC 360-16 and GB 50017-2017. Design considerations and differences of flexure buckling, torsion buckling and flexure-torsion buckling for two country's codes are introduced. These will be helpful to understand design accomplishments for three kinds of buckling.
The stability strength calculation of axial compression member is listed in Chapter E of AISC 360-16, in which the design compressive strength is φcPn; the compressive coefficient of drag is φc=0.9; the nominal compressive strength Pn shall be the lowest value obtained based on the applicable limit states of flexural buckling, torsional buckling, and flexural-torsional buckling.
For flexural buckling, the number of column curves in AISC 360-16 decreases from three to one. The flexural buckling strength of axial compression member is Pn=FcrAg, where there are Fcr=0.658Fy/FeFy or Fcr=0.877Fe, and Fe=((π2E)/((Lc/r)2)).
The column curve in AISC 360-16 can be divided into two segments, with the limit slenderness ratio of 137 (Fy=235 MPa) and the normalized slenderness ratio of λc=(Lc/πr)√Fy/E=1.5. In the elastic stage, the Euler's critical force is taken as the reference, and the reduction factor 0.877 is introduced to consider the influence of the geometric imperfections of the member, which is clear in concept and consistent with Euler's formula. In the inelastic stage, the nonlinear properties of the generalized Euler's formula are taken into account after the material strength exceeds the proportional limit, and the geometric imperfections and the residual stresses of the member are also considered.
GB 50017-2017 adopts four column curves, flexural buckling strength of which can be calculated by N/φAf ≤ 1.0. Considering the initial imperfections and the residual stresses of the member, stability factor φ is determined by the critical force of flexural buckling of axial compression member and is applicable in both elastic and inelastic zones. The slenderness ratios in both main axes are first calculated and then the larger slenderness ratio is substituted into φ=(1/2λn2)[(α2+α3λn+λn2)-√(α2+α3λn+λn2)2-4λn2] or i>φ=1-α1λn2 to calculate stability factor φ.
When analyzing torsional buckling and flexural-torsional buckling, torsional buckling may occur for biaxial symmetric cross-section. For cross-sections with single symmetric axis or without symmetric axis, flexural-torsional buckling may happen. The buckling strength of torsional buckling or flexural-torsional buckling is Pn=FcrAg, where there are Fcr=0.658Fy/FeFy or Fcr=0.877Fe, and Fe takes the critical stress of torsional buckling or flexural-torsional buckling.
Thus it can be seen, torsional buckling and flexural-torsional buckling of AISC 360-16 are calculated according to the column curve of flexural buckling after the computation of the corresponding elastic critical force.
The critical force of torsional buckling in GB 50017-2017 is Nz=(1/i02)(GIt+((π2EIω)/l2)), by which divided A is consistent with that in AISC 360-16.
In specific applications, GB 50017-2017 adopts the concept of equivalent slenderness ratio, and the slenderness ratio for torsional buckling is calculated by λz=√I0/(It/25.7+Iω/lω2) and considered as the slenderness ratio for flexural buckling. The stability factor φ is then computed based on the above method, which is the same as the method adopted in AISC 360-16. The difference is that the elastic segment of column curve in AISC 360-16 is determined by the product of Euler's critical load and reduction factor, and the inelastic segment is closer to the test curve, while GB 50017-2017 takes the regression curve from test data as the column curve.
GB 50017-2017 adopts (NEy-Nyz)(Nz-Nyz)=Nyz2(ys2/i02) to calculate the critical flexural-torsional buckling force Nyz. In AISC 360-16, supposing xo=0, flexural-torsional buckling equation of uniaxial symmetric cross-section is (Fe-Fey)(Fe-Fez)=Fe2(yo/[ro)2. Essentially, the above two equations are consistent. In other words, the calculation formulas of flexural-torsional buckling strength in GB 50017-2017 and AISC 360-16 are the same.
For axial compression member, both AISC 360-16 and GB 50017-2017 present design formulas for flexural buckling, torsional buckling, and flexural-torsional buckling. Both the two specifications determine the stability factor φ of three buckling modes based on column curve of flexural buckling. AISC 360-16 adopts one column curve and GB 50017-2017 uses four column curves, both of which consider the influences of geometric imperfection and residual stress and agree well with relevant tests in different countries. Torsional buckling and flexural-torsional buckling in AISC 360-16 and GB 50017-2017 come from the elastic stability equilibrium equation, which are transformed into the critical force of flexural buckling through equivalent critical force and designed by column curve of flexural buckling.
