Experimental and numerical determination of the chloride penetration in cracked concrete

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氧化铝陶瓷多粒子冲蚀磨损的数值模拟

氧化铝陶瓷多粒子冲蚀磨损的数值模拟∗胡彪;纪秀林;段慧;丁伟【摘要】采用LS⁃DYNA有限元分析软件建立多粒子冲蚀氧化铝陶瓷的有限元模型，运用LS⁃DYNA求解器对冲蚀过程进行仿真，通过观察靶材等效应力的分布分析冲蚀机制。

结果表明：靶材体积磨损率随着冲蚀角度的增大而增大，在冲蚀角度达到90°时，体积磨损率达到最大值，表现出典型的脆性材料的冲蚀特性；靶材的体积磨损率随着冲蚀速度的增大而增大，且具有良好的线性增长关系；靶材的体积磨损率整体上随着冲击粒子粒径的增大而增大，但在达到临界尺寸的一段时间内会随着粒径的增大而减小；靶材会吸收粒子的一部分动能转化为自己的内能，但随着粒子冲击结束而离开靶材表面，靶材表面形成微裂纹以及部分单元失效，因此靶材的能量随着单元的失效而减小。

%LS⁃DYNA was used to establish the finite element model of multi particles impacting on alumina ceramics, and the erosion process was simulated by using LS⁃DYNA solver� The erosion mechanism was analyzed by observing the distribution of Von Mises stress of the target� The results show that volume loss rate of the target is increased with increas⁃ing the impact angle, and volume loss rate reaches the maximum value at the impact angle of 90°, which exhibits erosio n characteristics of typical brittle materials� Volume loss rate of the target is increased with increasing the impact velocity, and they have good linear growth relationship� Volume loss rate of the target is increased with increasing the impact parti⁃cle size as a whole, but it is decreased within a period of time when the impact particle size reaches a critical size� The target absorbs part of the particles kinetic energy and transforms it intointernal energy, and when the particles leave the target surface,micro⁃crack and some elements failure are formed on the target surface, therefore, the target total energy is decreased with the failure of the elements.【期刊名称】《润滑与密封》【年(卷),期】2015(000)004【总页数】5页(P49-53)【关键词】冲蚀磨损;氧化铝陶瓷;脆性材料;多粒子;冲蚀机制【作者】胡彪;纪秀林;段慧;丁伟【作者单位】河海大学常州校区机电工程学院江苏常州213022;河海大学常州校区机电工程学院江苏常州213022;河海大学常州校区机电工程学院江苏常州213022;河海大学常州校区机电工程学院江苏常州213022【正文语种】中文【中图分类】TH117.1冲蚀磨损是固体颗粒随着高速流体对材料表面冲击造成的材料损坏，是一个动态的失效过程[1]。

关于针对承担项目情况的解释

HR Planning System Integration and Upgrading Research ofA Suzhou Institution承担项目情况：1）国家自然科学基金“镁合金板材大变形成形机制与过程模拟研究”，编号：50405014，经费：23万元，项目起止年月：2005.1～2007.12，负责人。

2）国家自然基金项目，“辊弯成形全流程动态模拟技术研究”，编号：50375095, 经费：24万元，起止年月：2004.1～2006.12，主要参加人。

3）国家自然基金重点项目，“材料智能化近终成形加工技术的若干基础问题”，编号：50634010, 经费：180万元，起止年月：2007.01～2010.12，主要参加人。

4）国家973计划前期研究项目“材料制备新方法探索及性能研究”，编号：2006CB708600，总经费：1094万元，起止年月：2006.12～2008.11，主要参加人。

5）国家863计划重点项目，“高强高韧镁合金及其应用技术研究”，编号：002AA331120，经费：340万元，起止年月：2002.6～2005.6，主要参加人。

6）国防科工委民口配套项目，“XXX轴承的研究”，编号：MKPT-05-268，经费：165万元，起止年月：2005.1～2006.12，主要参加人。

7）上海市创新科技支撑计划项目子课题，“薄带连铸带钢力学性能及表面裂纹形成机理研究”，编号：07DZ1103，经费：80万元，起止年月：2008.2-2009.12，负责人。

8）教育部新世纪优秀人才计划项目，“镁合金板材变形机理与成形性能的宏微观研究”，编号：NCET-07-0545，经费：50万元，起止年月：2008.1-2010.12，负责人。

9）上海市重点基础研究项目，“ERW焊管排辊成形理论与工艺设计方法研究”，编号：09JC1407000，经费：30万元，起止年月：2009.9～2011.8，负责人。

一种对流传质传热模拟方法:考虑土豆片中三种不同流体速度

一种对流传质传热模拟方法：考虑土豆片中三种不同流体速度王小勇;刘显茜【摘要】许多文献报道了土豆片干燥模拟方法.然而,其并没有清晰解释传热传质耦合机理.本文建立一个模型来描述土豆片中的温度演化和水分迁移.创新点是考虑到土豆片中存在三种不同流体速度以及认为材料物理变量之间相互影响.模拟证明计算结果与文献实验具有良好的拟合度.同时研究了不同的风温,风速以及空气相对湿度对干燥的影响.%Many simulation methods of drying potato slices were reported. However, this did not explain the coupled mechanism of heat and mass transfer clearly. A model is built to describe the temperature evolution and moisture immigrates of potato slices. The innovation point is considering that the coefficients will alter over other physical variables and the existence of three different fluid velocity in the porous medium. Then, a good fitting between literature experiment and calculation result can be observed. And, the effects of parameters such as temperature, air velocity and air relative humidity towards drying rate are researched.【期刊名称】《价值工程》【年(卷),期】2018(037)008【总页数】5页(P176-180)【关键词】耦合机制;流体速度;物理变量【作者】王小勇;刘显茜【作者单位】昆明理工大学机电工程学院,昆明650500;昆明理工大学机电工程学院,昆明650500【正文语种】中文【中图分类】TK1240 引言食品干燥是一个重要的工业生产部门。

SPH方法在流体晃动中的研究应用进展

SPH方法在流体晃动中的研究应用进展1概述光滑粒子流体动力学方法（Smoothed Particle Hydrodynamics，SPH）是近30多年来逐步发展起来的一种无网格方法，SPH法的基本思想是，用一系列任意分布的粒子来代替整个连续介质流体，并用粒子集合和插值核函数来估算空间函数及其导数，于是所有的力学变量都由这些粒子负载，积分方程则通过离散粒子的求和得到估值，N-S方程由原来同时含有时间和空间导数的偏微分方程转化为只含有时间导数的微分方程。

从原理上说，只要质点的数目足够多，就能精确地描述力学过程。

虽然在SPH方法中，解的精度也依赖于质点的排列，但它对点阵排列的要求远远低于网格的要求。

由于质点之间不存在网格关系，因此它可避免极度大变形时网格扭曲而造成的精度破坏等问题，并且也能较为方便的处理不同介质的交界面。

SPH法的主要优点如下：对流项直接通过粒子的运动来模拟，完全消除了自由界面上的数值发散问题，保证了自由液面追踪的清晰准确；完全不需要网格，不仅免去了生成网格的麻烦，SPH是一种纯Lagrange算法，能避免Euler描述中的欧拉网格与材料的界面问题，这些优点使得SPH法可以方便地模拟具有自由液面的大变形的流体流动问题[1,2]。

当然，SPH算法也有它的一些问题和不足之处，关于SPH算法应用中出现的问题，Swegle[3]作过详细研究，这些问题分别是张力的不稳定性、收敛性的缺乏和零能量模式。

针对这些问题，已经提出了相应的改进算法，产生了各种改进的SPH算法，如规则化光滑粒子动力学（RSPH）算法[4]、自适应光滑粒子流体动力学（ASPH）算法[1]、修正光滑粒子动力学（CSPH）算法[5]等，并广泛应用于各种研究分析中。

Randlesa[6]、Shaofan Li[7]、Monaghan[8]、Liu[9]等的综述对SPH算法近期的发展和应用做了系统的总结。

2SPH算法在流体晃动中的研究应用进展光滑粒子流体动力学（SPH）无网格方法作为一种创新方法出现来替代标准网格技术，是计算力学中出现最早的无网格粒子方法之一，由Lucy[10]、Gingold[11]同时提出，用来解决天体物理学问题。

EXPERIMENTAL AND NUMERICAL MODAL ANALYSIS OF A CONCRETE

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Experimental and numerical investigations on the influence o

can occur, e.g. either occasionally paired with an overload (mixed mode overload) or permanently in terms of a mixed mode block loading as a combination of normal and shear stresses. By means of this change, the lifetime is inﬂuenced as well. Only few investigations of such mixed mode loading change effects on the lifetime are described in the literature [e.g. 3–6]. Within the scope of this paper, ﬁrstly the inﬂuence of mixed mode loading changes in terms of mixed mode (Mode ICMode II) overloads and block loadings, which are interspersed into a Mode I baseline level loading, is experimentally investigated. Secondly, a detailed elastic– plastic ﬁnite element analysis of the fatigue crack growth after mixed mode overloads, which are interspersed into a Mode I baseline level loading, is presented in order to understand the mechanisms of the load interaction effects.

文献翻译-非线性动力学的实验和转子轴承系统支持的行为的数值研究

附录A英文原文Experimental and Numerica Studies on Nonlinear Dynam Behavior of Rotor System Supported by Ball BearingsBall bearings are important mechanical components in high-speed turbomachinery that is liable for severe vibration and noise due to the inherent nonlinearity of ball ing experiments and the numerical approach, the nonlinear dynamic behavior of a ﬂexible rotor supported by ball bearings is investigated in this paper. An experimental ball bearing-rotor test rig is presented in order to investigate the nonlinear dynamic performance of the rotor systems, as the speed is beyond the ﬁrst synchroresonance frequency. The ﬁnite element method and two-degree-of-freedom dynamic model of a ball bearing are employed for modeling the ﬂexible rotor s ystem. The discrete model of a shaft is built with the aid of the ﬁnite element technique, and the ball bearing model includes the nonlinear effects of the Hertzian contact force, bearing internal clearance, and so on.The nonlinear unbalance response is observed by experimental and numerical analysis.All of the predicted results are in good agreement with experimental data, thus validating the proposed model. Numerical and experimental results show that the resonance frequency is provoked when the speed is about twice the synchroresonance frequency, while the subharmonic resonance occurs due to the nonlinearity of ball bearings and causes severe vibration and strong noise. The results show that the effect of a ball bearing on the dynamic behavior is noticeable in optimum design and failure diagnosis of high-speed turbomachinery. [DOI: 10.1115/1.4000586]Keywords: ball bearing, rotor, experiment, nonlinear vibrationA.1 IntroductionBall bearings are one of the essential and important components in sophisticated turbomachinery such as rocket turbopumps, aircraft jet engines, and so on. Because of the requirement of acquiring higher performance in the design and operation of ballbearings-rotor systems, accurate predictions of vibration characteristics of the systems, especially in the high rotational speed condition, have become increasingly important.Inherent nonlinearity of ball bearings is due to Hertzian contact forces and the internal clearance between the ball and the ring.Many researchers have devoted themselves to investigating the dynamiccharacteristics associated with ball bearings. Gustafsson et al. [1] studied the vibrations due to the varying compliance of ball bearings. Saito [2] investigated the effect of radial clearance in an unbalanced Jeffcott rotor supported by ball bearings using the numerical harmonic balance technique. Aktürk et al. [2] used a three-degree-of-freedom system to explore the radial and axial vibrations of a rigid shaft supported by a pair of angular contact ball bearings. Liew et al. [4] summarized four different dynamic models of ball bearings, viz., two or ﬁve degrees of freedom, with or without ball centrifugal force, which could be applied to determine the vibration response of ball bearing-rotor systems. Bai and Xu [5] presented a general dynamic model to predict dynamic properties of rotor systems supported by ball bearings. De Mul et al. [6] presented a ﬁve-degree-of-freedom (5DOF) model for the calculation of the equilibrium and associated load distribution in ball bearings. Mevel and Guyader [7] described different routes to chaos by varying a control parameter. Jang and Jeong [8] proposed an excitation model of ball bearing waviness to investigate the bearing vibration. Then, considering the centrifugal force and gyroscopic moment of ball, they developed an analytical method to calculate the characteristics of the ball bearing under the effect of waviness in Ref. [9]. Tiwari et al. [10,11] employed a two-degree-of-freedom model to analyze the nonlinear behaviors and stability associated with the internal clearance of a ball bearing.Harsha [12-14], taking into account different sources of nonlin-earity, investigated the nonlinear dynamic behavior of ball bearing-rotor systems. Gupta et al. [15] studied the nonlinear dynamic response of an unbalanced horizontal ﬂexible rotor supported by a ball bearing. With the aid of the Floquet theory, Bai et al. [16] investigated the effects of axial preload on nonlinear dynamic characteristics of a ﬂexible rotor supported by angular contact ball bearings. Using the harmonic balance method, Sinou [17]performed a numerical analysis to investigate the nonlinear unbalance response of a ﬂexible rotor supported by ball bearings.In the abovementioned studies, main attention has been paid to the ball bearing modeling and the dynamic properties analysis according to simple bearing-rotor models. With theoretical analysis and experiment, Yamamoto et al. [18] studied a nonlinear forced oscillation at a major critical speed in a rotating shaft,which was supported by ball bearings with angular clearances.Ishida and Yamamoto [19] studied the forced oscillations of a rotating shaft with nonlinear spring characteristics and internal damping. They found that a self-excited oscillation appears in the wide range above the major critical speed. A dynamic model was derived, and experiments are carried out with a laboratory test rig for studying the misaligned effect of misaligned rotor-ball bearing systems in Ref. [20]. Tiwari et al. [21] presented an experimental analysis to study the effect of radial internal clearance of a ball bearing on the bearingstiffness of a rigid horizontal rotor. These experimental results validated theoretical results reported in their literatures [10,11]. Recently, Ishida et al. [22] investigated theoretically and experimentally the nonlinear forced vibrations and parametrically excited vibrations of an asymmetrical shaft supported by ball bearings. Mevel and Guyader [23] used an experimental test bench to conﬁrm the predicted routes to chaos in their previous paper [7]. It is noticeable of lack of experiments on nonlinear dynamic behavior of ﬂexible rotor systems supported by ball bearings. In Ref. [24], the ﬁnite element method was used to model a LH2 turbopump rotor system supported by ball bearings. Numerical results show that the subharmonic resonance, as well as synchroresonance, occurs in the start-up process. It is found that the subharmonic resonance is an important dynamic behavior and should be considered in engineering ball bearing-rotor system design. But, the experimental and numerical studies of the subharmonic resonance in ball bearing-rotor systems are very rare.With respect to the above, the present study is intended to cast light on the subharmonic resonance characteristics in ball bearing-rotor systems using experiments and numerical approach. An experiment on an offset-disk rotor supported by ball bearings is carried out, and the ﬁnite element method and two-degree-of-freedom model of a ball bearing are employed for modeling this rotor system. The predicted results are compared with the test data, and an investigation is conducted in the nonlinear dynamic behavior of the ball bearings-rotor system.2 Experimental InvestigationAn experimental rig is employed for studying the nonlinear dynamic behavior of ball bearing-rotor systems, as shown in Fig.1. The horizontal shaft is supported by two ball bearings at both ends, and the diskis mounted unsymmetrically. The shaft is coupled to a motor with a ﬂexible coupling. The motor speed is controlled with a feedback controller, which gets the signals from an eddy current probe. Four eddy current probes, whose resolution is 0.5 m, are mounted close to the disk and bearing at the right end in the horizontal and vertical directions, respectively. The displacement signals, obtained with the help of probes, are input into an oscilloscope to describe the motion orbit, and a data acquisition and processing system were used to analyze the effects of ball bearings on the nonlinear dynamic behavior. The data acquisition and processing system utilizes a full period sampling as the data acquisition method. Its sampling rate is 500 kHz maximum, and sample size is 12 bits. The system provides eight channels for vibratory response acquisition and 1 channel for rotational speed acquisition. All channels are simultaneous.The limitation with the presented experimental setup is that the maximum attainable speed is 12,000 rpm. The ﬁrst critical speed of the rotor system falls in the speed span, as the shaft is ﬂexible and its ﬁst synchroresonance frequency is near 66 Hz (3960rpm).Thus, the dynamic behavior can be studied as the speed is beyond twice the synchroresonance frequency.3 Rotor Dynamic ModelThe bearing-rotor system combines an offset-disk and two ball bearings, which support the rotor at both ends. The sketch map of the system is described in Fig. 2, where the frame oxyz is the inertial frame. The corresponding experiment assembly is shown in Fig. 3.3.1 Equations of Motion . Deﬁne ux and uy as the transverse deﬂections along the ox and oy directions, and x θ and y θ as the corresponding bending angles in the oxz and oyz planes, respectively. When x u 1, y u 1,x 1θ , and y 1θ denote the displacements of the ball bearing center location at the left end, the complex variables 1u and 1θ can be assumed asDenote the displacements of the disk center by 2u and 2θ, and the displacements of the ball bearing center location at the right end by 3u and 3θ. Using the ﬁnite element method, the equations of motion for the rotor system can be written as [25,26]where []M , []C , []K , and []G are the mass, damping, stiffness, and gyroscopic matrix of the rotor system, respectively, ω is the rotational speed, and {}u is the displacement vector{}g F and {}u F are the vectors of gravity load and unbalance forces.{}bF is the vector of nonlinear forces associated with ball bearings.3.2 Ball Bearing Forces. A ball bearing is depicted in a frame of axes oxyz in Fig.4. The contact deformation for the j-th rolling element j δis given aswhere i c and o c are the internal radial clearance between the inner,outer race, and rolling elements, respectively, in the direction of contact, and ubx and uby are the relative displacements of the inner and outer race along the x and y directions, respectively. As shown in Fig. 4, the angular location of the j-th rolling element j ϕ can be obtained fromWhere N , c ω, t , and 0ϕ are the number of rolling elements, cage angular velocity, time, and initial angular location, respectively. The cage angular velocity can be expressed as [27]where b D and p D are the ball diameter and bearing pitch diam- eter,respectively. α is the contact angle, which is concerned with the clearance and can be obtained as follows:Referring to Fig. 4, i r and o r are the inner and outer groove radius,respectively.If the contact deformation j δ is positive, the contact force could be calculated using the Hertzian contact theory; otherwise, no load is transmitted. The contact force j Q between the j-th ball and race can be expressed as follows:where b k is the contact stiffness that can be given bywhere bi k and bo k are the load-deﬂection constants between the inner and outer ball race, respectively[28]. Summing the contact forces for each rolling element, the total bearing reaction fb in a complex form is4 Experimental and Numerical AnalysisAs shown in Fig. 2, the experimental assembly and the ﬁnite element model used in the dynamic analysis represent the ball bearing-rotor system with the following geometrical properties:length between the disk center and left end bearing center mm L 1201=; length between the disk center and right end bearing center mm L 1202=; and the shaft diameter mm D 10=. In addition, the elastic shaft material is steel of density 37950m kg =ρ,Young’s modulus GPa E 211=, and Poisson’s ratio 3.0=v . The ball bearings at both ends are the same model, 7200AC, and its parameters are listed in Table 1.The unbalance load is acted wit h the aid of the mass ﬁxed on the disk. By virtue of this act, the mass eccentricity of the disk can be deﬁnitely ascertained. As the mass eccentricity of the disk is 0.032 mm, the vibratory response at different rotational speed is determined via a numerical integration and Newton –Raphson iterations of the nonlinear differential equation (2). Note that the clearances used to simulate the bearings are measured ones. The horizontal and vertical displacements signals near the disk are acquired at different times, along with the increased rotational speed. Thus, the amplitudes of vibration at different speeds are determined according to the test data, and overall amplitudes are illustrated in Fig. 5, as the rotor system is run from 2000=ω rpm to 10,000 rpm. The prediction results compared with experimental data are shown in Fig. 5. It can be found that all of the predicted results are in good agreement with experimental data, thus validating the proposed model. The ﬁrst predicted resonance peak—the so called forward critical speed in linear theory,located at3960=ω rpm, matches the experimental date near 3960=ω rpm quite well. Moreover, the other amplitude peak appearing in the rotational speedrange7700=ω rpm to 8100 rpm can be found in both experimental and numerical analysis results.The corresponding frequency value of this peak is just the frequency doubling of the system critical speed.The Floquet theory can be used for analyzing the stability and topological properties of the periodic solution of the ball bearingrotor system. If the gained Floquet multipliers are less than unity,the periodic solution of the system is stable. If at least one Floquet multiplier exists with the absolute value higher than unity, the periodic solution is unstable and the topological properties of response alter into nonperiodic motion [29]. The leading Floquet multipliers and its absolute value at 7600=ω rpm, 8029 rpm, and 8200 rpm are shown in Table 2. It is found that the leading Floquet multiplier of the system remains in the unit circle, which indicates a synchronous response, as the rotational speed is less than 7700 rpm. Stability analysis shows that the imaginary part of the two leading Floquet multipliers move in opposite directions along the real axis near 7700=ω rpm. When the speed exceeded 7700=ω rpm, the leading Floquet multiplier crosses the unit circle through -1, as shown in Table 2. The periodic solution loses stability and undergoes a period-doubling bifurcation to a period-2 response, which indicates that a subharmonicresonance occurs. The subharmonic resonance keeps on from 7700=ω rpm to 8100 rpm. At 8100=ω rpm, the leading Floquet multiplier moves inside the unit circle through -1. Imply that the subharmonic resonance vanishes and the synchronous response returns. The synchronous response then continues to exist forspeeds above 8100=ω rpm.The waterfall map of frequency spectrums comparisons for prediction and experiment results are illustrated in Fig. 6. It can be found that agreement between the prediction and the experimental data is remarkable. The frequency component 66.9 Hz, near the forward resonance frequency, emerges and its amplitude rises signiﬁcant when the rotational speed is near 8029 rpm. It is shown that the resonance frequency is provoked when the speed is about twice the critical speed of the ball bearing-rotor system, and the subharmonic resonance occurs. The experimental and numerical analysis indicate that the representative nonlinear behavior and the subharmonic resonance arise from the nonlinearity of ball bearings, Hertzian contact forces, and internal clearance.The orbit and frequency spectrum at 8029=ω rpm are plotted in Fig. 7. Not only the prediction orbit but also the experiment results imply that the response is a period-2 motion, which is illustrated in Fig. 7(a). The predicted frequency components, consisting of 8.133=ω Hz (8029 rpm) and 9.662=ω Hz (4014rpm), coincide with experimental data. It indicates that the periodic response loses stability through a period-doubling bifurcation to a period-2 response. Thus, the subharmonic resonance occurs due to the effects of ball bearings. It can cause severe vibration and strong noise. Moreover, the subharmonic resonance could couple with other destabilizing effects on engineering rotor systems such as Alford forces, internal damping, and so on, and induce the rotor to lose stability and damage.5 ConclusionsAn experimental rig is employed to investigate the nonlinear dynamic behavior of ball bearing-rotor systems. The corresponding dynamic model is established wi th the ﬁnite element method and 2DOF dynamic model of a ball bearing, which includes the nonlinear effects of the Hertzian contact force and bearing internal clearance. All of the predicted results are in good agreement with experimental data, thus validating the proposed model. Numerical and experimental results show that the resonance frequency is provoked, and the subharmonic resonance occurs due to the nonlinearity of ball bearings when the speed is about twice the synchroresonance frequency. The subharmonic resonance cannot only cause severe vibration and strong noise, but also induce the rotor to lose stability and damage, once coupled with other destabilizing effects on high-speed turbomachinery such as Alford forces, internal damping, and so on. It is found that the effect of the Hertzian contact forces could also induce a subharmonic resonance, even if the internal clearance was not present. But, the response amplitude and subharmonic component of the rotor system without internal clearance are less than that with both Hertzian contact forces and internal clearance. Otherwise, the clearance may be unavoidable under high-speed operations, where the bearings are axially preloaded since the effect of unbalanced load is signiﬁcant at high speed. Thus, the nonlinearity of ball bearings,Hertzian contact forces, and internal clearance should be taken into account in ball bearing-rotor system design and failure diagnosis.AcknowledgmentThe authors would like to acknowledgment the support of the National Natural Science Foundation of China (Grant No.10902080) and Natural Science Foundation of Shaanxi Province(Grant Nos. SJ08A19 and 2009JQ1008).References[1] Gutafsson, O., and Tallian, T., 1963, “Resear ch Report on Study of the Vibration Characteristics of Bearings,” SKF Ind. Inc. Technical Report No.AL631023.[2] Saito, S., 1985, “Calculation of Non-Linear Unbalance Response of Horizontal Jeffcott Rotors Supported by Ball Bearings With Radial Clearances,” ASME J.Vib., Acou st., Stress, Reliab. Des., 107(4), pp. 416–420.[3] Aktürk, N., Uneeb, M., and Gohar, R., 1997, “The Effects of Number of Balls and Preload on Vibrations Associated With Ball Bearings,” ASME J. Tribol.,119, pp. 747–753.[4] Liew, A., Feng, N., and Hahn, E., 2002, “Transient Rotordynamic Modeling of Rolling Element Bearing Systems,” ASME J. Eng. Gas Turbines Power,124(4), pp. 984–991.[5] Bai, C. Q., and Xu, Q. Y., 2006, “Dynamic Model of Ball Bearing With Internal Clearance and Waviness,” J. Sound Vib., 294(1-2), pp. 23–48.[6] De Mul, J. M., Vree, J. M., and Maas, D. A., 1989, “Equilibrium and Associated Load Distribution in Ball and Roller Bearings Loaded in Five Degrees of Freedom While Neglecting Friction—Part I: General Theory and Application to Ball Be arings,” ASME J. Tribol., 111, pp. 142–148.[7] Mevel, B., and Guyader, J. L., 1993, “Routes to Chaos in Ball Bearings,” J.Sound Vib., 162, pp. 471–487.[8] Jang, G. H., and Jeong, S. W., 2002, “Nonlinear Excitation Model of Ball Bearing Waviness in a Rigid Rotor Supported by Two or More Ball Bearings Considering Five Degrees of Freedom,” ASME J. Tribol., 124, pp. 82–90.[9] Jang, G. H., and Jeong, S. W., 2003, “Analysis of a Ball Bearing With Waviness Considering the Centrifugal Force and Gyroscopic Moment of the Ball,”ASME J. Tribol., 125, pp. 487–498.[10] Tiwari, M., Gupta, K., and Prakash, O., 2000, “Effect of Radial Internal Clearance of a Ball Bearing on the Dynamics of a Balanced Horizontal Rotor,” J.Sound Vib., 238(5), pp. 723–756.[11] Tiwari, M., Gupta, K., and Prakash, O., 2000, “Dynamic Response of an Unbalanced Rotor Supported on Ball Bearings,” J. Sound Vib., 238(5), pp.757–779.[12] Harsha, S. P., 2005, “Non-Linear Dynamic Response of a Balanced Rotor Supported on Rolling Element Bearings,” Me ch. Syst. Signal Process., 19(3),pp. 551–578.[13] Harsha, S. P., 2006, “Rolling Bearing Vibrations—The Effects of Surface Waviness and Radial Internal Clearance,” Int. J. Computational Methods in Eng Sci. and Mech., 7(2), pp. 91–111.[14] Harsha, S. P., 2006, “Nonlinear Dynamic Analysis of a High-Speed Rotor Supported by Rolling Element Bearings,” J. Sound Vib., 290(1–2), pp. 65–100.[15] Gupta, T. C., Gupta, K., and Sehqal, D. K., 2008, “Nonlinear Vibration Analysis of an Unbalanced Flexible Rotor Supported by Ball Bearings With Radial Internal Clearance,” Proceedings of the ASME Turbo Expo, Vol. 5, pp. 1289–1298.[16] Bai, C. Q., Zhang, H. Y., and Xu, Q. Y., 2008, “Effects of Axial Preload of Ball Bearing on theNonlinear Dynamic Characteristics of a Rotor-Bearing System,” Nonlinear Dyn., 53(3), pp. 173–190. [17] Sinou, J. J., 2009, “Non-Linear Dynamics and Contacts of an Unbalanced Flexible Rotor Supported on Ball Bearings,” Mech. Mach. Theory, 44(9), pp.1713–1732.[18] Yamamoto, T., Ishida, Y., and Ikeda, T., 1984, “Vibrations of a Rotating Shaft With Rotating Nonlinear Restoring Forces at the Major Critical Speed,” Bull.JSME, 27(230), pp. 1728–1736.[19] Ishida, Y., and Yamamoto, T., 1993, “Forced Oscillations of a Rotating Shaft With Nonlinear Spring Characteristics and Internal Damping (1/2 Order Subharmonic Oscillations and Entrainment),” Nonlinear Dyn., 4(5), pp. 413–431.[20] Lee, Y. S., and Lee, C. W., 1999, “Modeling and Vibration Analysis of Misaligned Rotor-Ball Bearing Systems,” J. Sound Vib., 224(1), pp. 17–32.[21] Tiwari, M., Gupta, K., and Prakash, O., 2002, “Experimental Study of a Rotor Supported by Deep Groove Ball Bearing,” Int. J. Rotating Mach., 8(4), pp.243–258.[22] Ishida, Y., Liu, J., Inoue, T., and Suzuki, A., 2008, “Vibrations of an Asymmetrical Shaft With Gravity and Nonlinear Spring Characteristics (IsolatedResonances and Internal Resonances),” ASME J. Vib. Acoust., 130(4),p.041004.[23] Mevel, B., and Guyader, J. L., 2008, “Experiments on Routes to Chaos in Ball Bearings,” J. S ound Vib., 318, pp. 549–564.[24] Bai, C. Q., Xu, Q. Y., and Zhang, X. L., 2006, “Dynamic Properties Analysis of Ball Bearings—Liquid Hydrogen Turbopump Used in Rocket Engine,”ACTA Aeronaut. Astronaut. Sinica, 27(2), pp. 258–261. [25] Nelson, H., 1980, “A Finite Rotating Shaft Element Using Timoshenko Beam Theory,” ASME J. Mech. Des., 102(4), pp. 793–803.[26] Zhang, W., 1999, Basis of Rotordynamic Theory, Science Press, Beijing,China, Chap. 3.[27] Harris, T. A., 1984, Rolling Bearing Analysis, 2nd ed., Wiley, New York.[28] Aktürk, N., 1993, “Dynamics of a Rigid Shaft Supported by Angular Contact Ball Bearings,” Ph.D. thesis, Imperial College of Science, Technology and Medicine, London, UK.[29] Zhou, J. Q., and Zhu, Y. Y., 1998, Nonlinear Vibrations, Xi’an Jioatong University Press, Xi’an, China.附录B英文翻译非线性动力学的实验和转子轴承系统支持的行为的数值研究深沟球轴承在高速流体机械部件承担严重的振动和噪声的固有的非线性是很重要的。

注浆加固路基施工方法探讨

Roads and Bridges 道路桥梁53注浆加固路基施工方法探讨王子玉1,2 田立慧3（1.海南热带海洋学院生态环境学院，海南三亚 572022）（2.哈尔滨工业大学土木工程学院，黑龙江哈尔滨 150090）（3.黑龙江科技大学矿业工程学院，哈尔滨 150090）中图分类号：U45 文献标识码：B 文章编号1007-6344（2019）06-0053-01摘要：采用黏土浆液材料注入路基中，使路基具有显著抗冻性和隔水效果，能够在路基内部形成有效防渗帷幕，切断地下水渗流通道，从根本上解决季节性冻融路基由于冻融期间水份迁而引起的不均匀冻胀问题。

黏土浆液材料造价低廉，直接针对现行普通水泥浆液和水泥黏土浆液性能上的某些不足，具有多项优于现行普通水泥浆液和水泥黏土浆液的良好性能。

关键词：季节冻土区；铁路路基；注浆法加固；施工方法0 引言化学加固不同于其它的工程加固措施，它着眼于提高土体本身的工程性质，依据原土的物理及化学构成和工程条件的不同，可以选用不同的加固材料和施工工艺，使加固材料和原土发生完全或部分的混合，产生一系列复杂的物理、化学反应，包括离子交换、火山灰反应、凝聚作用和硬化促进反应等。

从而使土体的承载力、密实度和耐水性都得到较大提高。

化学加固除浅层地表土在填筑过程中可以使用拌合法外，对于基础工程，可依加固材料和施工工艺的不同分为粉体喷射搅拌法和注浆法两大类。

近年来路基冻害的防治研究一直进行着，王彦虎等采用钻孔埋管注盐法整治路基冻害，冻胀量下降明显[1]。

汪双杰，陈建兵，刘华，牛富俊，许健，吕菲，闫宏业等采用保温法抑制路基冻胀效果，并对保温材料厚度、路基填料等进行研究[2-7]。

金兰、冷景岩挺等通过改善路基填料的技术方法来提高防冻胀需求[8-9]。

杨有海等提出在边坡及护道表面加以封闭，并设置隔水墙、排水沟、渗水沟、蒸发池等，使路基水分得以控制，进而治理路基冻害[10]。

1 注浆加固施工方法该注浆法使用传统的注浆设备，施工简单易行。

线热源和柱热源

现在通常用到的地埋管换热器（BHE）设计理论主要有线热源理论和柱热源理论，线热源理论又分为无限长线热源理论和有限长线热源理论。

经典的无限长线热源理论最早是由Kelvin提出的[19]，之后Ingersoll等在该理论的基础上建立了忽略轴向传热的径向传热模型[20]。

该模型将BHE周围的岩土体视为具有初始均匀温度的无限介质，假定钻孔为无限线热源，忽略热源的末端效应[21]。

有限长线热源理论假设BHE中传热流体温度恒定，根据BHE的长度来对g函数进行求解，Eskilson是第一个给出g函数解析式的人[22]。

线热源理论物理意义明确，应用简单，对持续恒热流运行的BHE有很好的解析精度，但是当BHE运行小于6h时会产生明显错误，此时，需要使用柱热源理论来进行分析计算[20]。

柱热源理论是由Carslaw和Jaeger首次提出的，类似于无限线热源理论，柱热源理论也假设圆柱体被无限均匀介质包围[23]。

柱热源理论提出后，Ingersoll，Kavanaugh和Hellstrom等都对该理论的修改完善做出了重要贡献[7,20,24]。

上个世纪90年代，我国有关BHE设计理论的研究才真正开始。

方肇洪、刁乃仁、曾和义等将线热源理论和柱热源理论引入国内，并在此基础上进行了大量的研究，建立了BHE二维或三维的数学模型[8,25-27]。

李新国考虑土壤类型、热物性及湿度迁移等各方面因素，将埋设于土壤中的BHE处理为等效内热源，提出内热源型BHE理论模型，研究建立内热源型BHE周围土壤热湿传递物理和数学模型[28]。

於仲义将地埋管与周围土壤传热区域分为钻井内外两部分，钻井内为稳态传热，采用多极理论U型管模型，钻井外为瞬态传热，采用柱热源模型，并考虑了地埋管内循环介质轴向变化。

在一定的换热时间下，利用该模型分析地埋管传热特性，具有较高的可靠性和较少的计算工作量[29]。

虽然通过前期研究已初步建立了BHE设计理论体系，但是，该体系并不完善，主要是因为它们都是建立在持续恒热流基础之上。

激光一次性去除铝合金飞机蒙皮涂层实验与数值研究

表面技术第53卷第9期激光一次性去除铝合金飞机蒙皮涂层实验与数值研究姜苏航，李多生*，叶寅，谢非彤，邱彦钦，钟宏平（南昌航空大学，南昌 330063）摘要：目的研究激光的频率、功率及扫描速度等参数对脉冲激光清洗航空铝合金表面S06-0215油漆涂层的影响，分析脉冲激光清洗的机制，优化工艺参数组合，并设计航空铝合金表面涂层一次性去除方法。

方法以航空铝合金2A12为基材，开展脉冲激光的频率、功率和扫描速度等参数对基材表面涂层烧蚀深度的影响研究，以及烧蚀过程中基材表面最高温度的模拟研究。

同时，以表面粗糙度和去除深度为评价指标，对2A12铝合金飞机蒙皮表面涂层进行清洗实验，对采用优化参数清洗后的蒙皮表面进行粗糙度测量、元素含量以及组成成分分析。

结果脉冲激光的扫描速度和频率变大，以烧蚀为主导作用的去除机制逐渐减弱，同时振动机制逐渐增强，并最终转变为主导地位。

影响激光去除深度的参数，按权重大小依次为扫描速度、功率、频率。

结合模拟与实验结果发现，激光频率125 kHz、功率70 W和速度50 mm/s为表面涂层S06-0215最佳的一次性去除工艺参数组合，此时能量密度大小为1.47 J/cm2，清洗过程中，不损伤2A12铝合金飞机蒙皮基材。

通过XRD、SEM以及EDS表征分析表明，氧化膜仅被部分去除。

结论激光清洗铝合金表面油漆涂层采用合适的参数组合可以保留氧化膜，可以实现一次性地去除铝合金表面油漆涂层，同时对基材不会造成损伤。

关键词：激光清洗；纳秒脉冲激光；铝合金；飞机蒙皮；一次性去除；正交实验法中图分类号：TG174 文献标志码：A 文章编号：1001-3660(2024)09-0180-10DOI：10.16490/ki.issn.1001-3660.2024.09.017Experimental and Numerical Study on Laser One-time Removal ofAluminium Alloy Aircraft Skin CoatingJIANG Suhang, LI Duosheng*, YE Yin, XIE Feitong, QIU Yanqin, ZHONG Hongping(Nanchang Hangkong University, Nanchang 330063, China)ABSTRACT: The work aims to study the effect of three parameters such as frequency, power and scanning speed of the pulsed laser on cleaning of the paint coating (S06-0215) on the surface of aluminum alloy aircraft skin in detail, then analyze the mechanism of pulsed laser cleaning, and finally optimize the combination of process parameters and designed one-time removal method for removing the paint coating of aircraft skin. Aerospace aluminum alloy (2A12) was used as the substrate, and the收稿日期：2023-02-20；修订日期：2023-09-08Received：2023-02-20；Revised：2023-09-08基金项目：国家自然科学基金（51562027，12062016，51975287）；江西省重点研发计划重点项目（20201BBE51001）；江西省省级优势科技创新重点团队项目（20181BCB24007）；江苏省重点研发计划（BE2021055）Fund：The National Natural Science Foundation of China (51562027, 12062016, 51975287); The Key Project of Jiangxi Province Key R&D Program (20201BBE51001); The Jiangxi Province Provincial Advantageous Science and Technology Innovation Key Team Project (20181BCB24007); The Key R&D Plan of Jiangsu Province (BE2021055)引文格式：姜苏航, 李多生, 叶寅, 等. 激光一次性去除铝合金飞机蒙皮涂层实验与数值研究[J]. 表面技术, 2024, 53(9): 180-189.JIANG Suhang, LI Duosheng, YE Yin, et al. Experimental and Numerical Study on Laser One-time Removal of Aluminium Alloy Aircraft Skin Coating[J]. Surface Technology, 2024, 53(9): 180-189.*通信作者（Corresponding author）第53卷第9期姜苏航，等：激光一次性去除铝合金飞机蒙皮涂层实验与数值研究·181·frequency, power and scanning speed of pulsed laser were optimized by orthogonal design. The optimized process parameters were applied to simulate the final ablation depth of the substrate surface coating and the maximum temperature of the substrate surface during the ablation process, and then the surface roughness and removal depth after cleaning were used as evaluation criteria to judge the effect on cleaning of the aluminum alloy aircraft skin coating. The orthogonal experiments were carried out to optimize the removal parameter combination, and the final combination of parameters was used to measure the roughness of the cleaned skin surface. Next, the microscopic morphology and element content were analyzed. The simulation and experimental results showed that the ablation was dominant effect, which was the main removal mechanism. With scanning speed and frequency increasing, the removal mechanism dominated by ablation gradually weakened, while the vibration mechanism gradually increased and eventually became dominant. The energy density of the pulsed laser was determined by the power and frequency, and it was proportional to the power and inversely proportional to the frequency. The energy density of the laser and the scanning speed together determined the quality of the cleaned surface. When energy density was small, the depth of laser cleaning was shallow. With energy density increasing, the thermal impact of the laser range was larger, and the depth of cleaning was also deeper. The frequency of the laser had a certain effect on the surface roughness after cleaning. The greater the frequency, the greater the number of laser actions per unit of time and the more obvious the vibration effect. The weight order of parameters affecting laser removal depth was as follows: scanning speed > laser power > pulse frequency. It was found that when laser frequency was 125 kHz, power was 70 W and speed was 50 mm/s, it was the best one-time removal process parameter combination for cleaning the surface coating (S06-0215), namely that energy density was 1.47 J/cm2 after cleaning. Under this parameter, not only the skin surface coating could be completely removed, but also the substrate surface oxide film could be completely retained to achieve the non-destructive cleaning effect. At that time, no special elements of coating and substrate were found on the micro-surface after cleaning, and the macro-surface roughness was lower than that before cleaning. Therefore, this study effectively combines the optimized parameters of orthogonal experiment with the predicted results of simulation, and innovatively develops a novel efficient laser cleaning method to remove the coating of aircraft skin at one-time, which provides an important reference for the development of new efficient and accurate laser cleaning of multi-layer coating.KEY WORDS: laser cleaning; nanosecond pulsed laser; aluminum alloy; aircraft skin; one-time removal; orthogonal experimental method近年来，随着激光技术的快速发展，激光产业领域也越来越大[1-3]。

湍流长度尺度英文

湍流长度尺度英文Turbulence Length ScalesTurbulence is a complex and fascinating phenomenon that has been the subject of extensive research and study in the field of fluid mechanics. One of the key aspects of turbulence is the concept of turbulence length scales, which refers to the range of different-sized eddies or vortices that are present in a turbulent flow. These length scales play a crucial role in understanding and predicting the behavior of turbulent flows, and they have important implications in a wide range of engineering and scientific applications.The smallest length scale in a turbulent flow is known as the Kolmogorov length scale, named after the Russian mathematician and physicist Andrey Kolmogorov. This length scale represents the size of the smallest eddies or vortices in the flow, and it is determined by the rate of energy dissipation and the kinematic viscosity of the fluid. The Kolmogorov length scale is typically denoted by the Greek letter η (eta) and can be expressed as η = (ν^3/ε)^(1/4), where ν is the kinematic viscosity of the fluid and ε is the rate of energy dissipation.The Kolmogorov length scale is important because it represents the scale at which viscous forces become dominant and energy is dissipated into heat. Below this length scale, the flow is considered to be in the dissipation range, where the eddies are too small to sustain their own motion and are rapidly broken down by viscous forces. The Kolmogorov length scale is therefore a critical parameter in the study of turbulence, as it helps to define the range of scales over which energy is transferred and dissipated within the flow.Another important length scale in turbulence is the integral length scale, which represents the size of the largest eddies or vortices in the flow. The integral length scale is typically denoted by the symbol L and is a measure of the size of the energy-containing eddies, which are responsible for the bulk of the turbulent kinetic energy in the flow. The integral length scale is often determined by the geometry of the flow domain or the boundary conditions, and it can be used to estimate the overall scale of the turbulent motion.Between the Kolmogorov length scale and the integral length scale, there is a range of intermediate length scales known as the inertial subrange. This range is characterized by the presence of eddies that are large enough to be unaffected by viscous forces, but small enough to be unaffected by the large-scale features of the flow. In this inertial subrange, the energy is transferred from the large eddies to the smaller eddies through a process known as the energycascade, where energy is transferred from larger scales to smaller scales without significant dissipation.The energy cascade is a fundamental concept in turbulence theory and is described by Kolmogorov's famous 1941 theory, which predicts that the energy spectrum in the inertial subrange should follow a power law with a slope of -5/3. This power law relationship has been extensively verified through experimental and numerical studies, and it has important implications for the modeling and prediction of turbulent flows.In addition to the Kolmogorov and integral length scales, there are other important length scales in turbulence that are relevant to specific applications or flow regimes. For example, in wall-bounded flows, the viscous length scale and the boundary layer thickness are important parameters that can influence the turbulent structure and behavior. In compressible flows, the Taylor microscale and the Corrsin scale are also relevant length scales that can provide insight into the characteristics of the turbulence.The understanding of turbulence length scales is crucial for a wide range of engineering and scientific applications, including fluid dynamics, aerodynamics, meteorology, oceanography, and astrophysics. By understanding the different length scales and their relationships, researchers and engineers can better predict andmodel the behavior of turbulent flows, leading to improved designs, more accurate simulations, and a deeper understanding of the fundamental principles of fluid mechanics.In conclusion, turbulence length scales are a fundamental concept in the study of turbulent flows, and they play a crucial role in our understanding and modeling of this complex and fascinating phenomenon. From the Kolmogorov length scale to the integral length scale and the inertial subrange, these length scales provide valuable insights into the structure and dynamics of turbulence, and they continue to be an active area of research and exploration in the field of fluid mechanics.。

汽车碰撞事故中下肢的损伤容限与机制

CN 11-5904/U J Automotive Safety and Energy, 2010, Vol. 1 No. 4253—259汽车碰撞事故中下肢的损伤容限与机制陈海斌1，王正国1，Albert I King2，Liying ZHANG2（1. 第三军医大学大坪医院野战外科研究所，创伤、烧伤与复合伤国家重点实验室，重庆 400042;2. Bioengineering Center, Wayne State University, Detroit, Michigan 48202, USA）摘要：为开展汽车碰撞事故的实验分析和数值建模，综述了下肢的损伤容限数据和损伤机制。

被选文献起1859年的自股骨三点弯曲试验数据（Weber），直至2009年的行人下肢多刚体建模数据（Kerrigan）。

数据表明：在汽车碰撞事故中下肢损伤较为多见，其症状常常为大面积软组织撕裂或缺损，并伴有严重骨折或脱位。

对于股骨、髌骨、膝关节、胫骨和踝关节等下肢部位，比较了在静态、动态条件下的弯矩、扭矩、轴向压缩力的损伤容限数据。

对于膝关节伤、长骨干骨折、股骨颈骨折、股骨髁骨折、踝关节骨折、脚骨骨折等典型伤类，讨论了下肢撞击的损伤机制。

关键词：交通事故；碰撞；下肢；损伤；容限；生物力学中图分类号：Q66Injury tolerance and mechanism of lower extremity inautomobile impact accidentsCHEN Haibin1, WANG Zhengguo1, Albert I King2, Liying ZHANG2(1. State Key Laboratory of Trauma, Burns, and Combined Injuries, Institute of Surgery Research,Daping Hospital, Third Military Medical University, Chongqing 400042, China;2. Bioengineering Center, Wayne State University, Detroit, Michigan 48202, USA)Abstract: A review about injury tolerance and injury mechanism of lower extremity is given to conduct the experimental studiesand numerical modeling for automobile impact accidents. Data sources were selected from Weber (1859, cadaver femur three-point bending tests) to as recent as Kerrigan (2009, multibody modeling of pedestrian lower extremity). One epidemiologicalinvestigation was described where lower extremity injuries are found to be the common form of injury associated with automobileimpact accidents, generally with massive soft tissue tear or defect and severe bone fracture or dislocation. Injury tolerance of thecommonly-injured regions of the lower extremities, including femur, patella, knee, tibia, and ankle, was depicted primarily in termsof the peak axial compressive force or bending/torsional moment for static and dynamic conditions. The injury mechanism offollowing injury patterns is summarized including knee joint injury, long bone shaft fracture, femoral neck fracture, femoral condylefracture, ankle joint injury, and foot bone fracture.Key words: traffic accidents; impact; lower extremity; wounds and injuries; threshold limit values; biomechanics收稿日期/ Received：2010-09-07基金项目/ Supported by：国家自然科学基金海外及港澳学者合作研究项目（30928005）；GM-国家自然科学基金（30122202）；重庆市自然科学基金（CSTC2009BB5013）资助项目第一作者/ First author：陈海斌/ CHEN Haibin（1965—），男（汉），湖北，副研究员。

热-力耦合作用下复合材料的跨尺度分析-热力学论文-物理论文

热-力耦合作用下复合材料的跨尺度分析-热力学论文-物理论文——文章均为WORD文档，下载后可直接编辑使用亦可打印——材料热力学论文教授推荐8篇之第五篇：热-力耦合作用下复合材料的跨尺度分析摘要：由于复合材料内部纤维与树脂的热膨胀系数差异很大，尤其是树脂性能对温度载荷较为敏感，服役时复合材料环境的高低温变化将使其热力学性能与常温状态产生较大差异。

采用Maxwell本构模型，探讨了温度变化对树脂材料本构关系的影响。

假设纤维为稳定材料，即其性能不随温度变化，依据复合材料细观力学理论选择六边形代表体积元为分析对象，建立了复合材料在温度载荷下热力学的本构模型。

并分别讨论了温度载荷下复合材料内部纤维体分比和纤维排列方式变化对其热力学性能的影响，实现了热-力耦合作用下复合材料的跨尺度分析。

关键词：复合材料；本构；跨尺度；细观力学有限元；Abstract:As the temperature of severing environment variating severely,there would be a great alteration in the thermodynamics properties of composites. Due to the great difference of thermal expansion coefficient between internal fiber and resin in the composites,the resin performance is more sensitive to temperature environment. In this paper,based on the Maxwell constitutive model,the relationship between resin constitutive and temperature is studied. Assuming that the fiber performance does not vary with temperature,according to the mesoscopic composite theory,a hexagon representative volume element is employed,and the finite element method（FEM）is adopted to establish the thermodynamics constitutive model of composites under temperature. The influence of fiber arrangement and fiber content on the composites thermodynamics constitutive is discussed,respectively. The process would be applied for a multi-scale solution for composite under thermodynamics coupling loads.Keyword:composite; constitutive; multi-scale; micromechanics FEM;1 前言具有轻质高强、耐腐蚀、抗疲劳等优良特性的复合材料日益受到低温工程的青睐，随着应用范围的扩大，对其温度载荷下的热力学性能响应的研究也逐渐深入和广泛。

Experimental and numerical analysis of the

Although in many engineering applications the design of structural parts is restricted to the elastic response of the materials involved, the knowledge of their behaviour beyond the elastic limit is relevant since plastic e ects with usually large deformations take place in the previous manufacturing procedures such as metal forming (traditionally classiÿed into (a) bulk-forming operations: forging, drawing, extrusion and rolling and (b) sheet forming operations: deep drawing, magnetic-pulse forming, tube enlarging and bulging [1]). Other important applications of elastoplastic models for metals are: crashworthiness, impact problems, inelastic buckling of thin-walled structures, superplastic forming, etc. (see, e.g. [2–4]).
In recent years, several ÿnite element large strain formulations usually deÿned within the plasticity framework have been developed and applied to the analysis of this test under isothermal and non-isothermal conditions (see, e.g. [4,6–10]). Moreover, some of such formulations have been validated, generally under isothermal conditions, with experimental data considering cylindrical specimens of di erent materials. In contrast, there have been only few studies focused on the necking phenomenon of strips (see, e.g. [11,12]).

平板太阳能集热器研究进展

平板太阳能集热器研究进展发布时间：2021-06-23T17:22:03.867Z 来源：《基层建设》2021年第8期作者：张宝喜1 赵丹2 王成顺1 韩铭泽1 [导读]1.沈阳建筑大学市政与环境工程学院沈阳 110168；2.沈阳市华域建筑设计有限公司沈阳 110168太阳能具有清洁、环保和普遍等优点，且有很多种利用方式，为了降低传统能源（如煤炭、石油和天然气等化石类不可再生能源）在建筑用能中的占比，将太阳能应用到建筑领域得到了国家的大力提倡，太阳能集热器就是将光转化为电或热的重要装置。

研究人员早在17世纪后期就已经研究出平板太阳能集热器，平板太阳能集热器是最早的太阳能热利用设备之一，但由于一些原因它在20世纪60年代以后才逐渐被深入研究和实际应用。

近年来，随着全球能源量的日趋紧张、人们节能环保意识的提高和国家政策的推动，平板太阳能集热器由于价格低廉和寿命长等优点逐渐得到了越来越多的开发和应用。

目前，平板太阳能集热器研究工作的主要任务是实现集热器结构优化、减少集热器热损失和提高集热器效率等。

国内外研究学者在此进行了大量的理论研究、实验研究和模拟研究，并取得了大量的成果。

1.平板太阳能集热器国外研究现状早在19世纪后期，平板太阳能集热器在美国西南部农场就有应用，且在1891年美国工程师Clarence Kemp发明了世界上第一台采用平板太阳能集热器的太阳能热水器[1]，1980年后，由于计算机技术的快速发展，理论研究在能够应用计算机进行模拟基础上发展迅速，通过分析平板太阳能集热器的流动换热原理和传热过程，形成了比较完善的理论体系。

2010年，Alvarez A[2]和Damir Dovic等[3]相继对平板太阳能集热器通过使用波纹型吸热板来改善传热效率的方法进行了研究。

两者的实验和模拟研究均表明波纹型吸热板在平板太阳能集热器中的应用可以达到提高集热器瞬时效率的目的。

2011年，Hanane D等[4]将摩洛哥气象参数作为参考，通过模拟研究了处于该地区的平板太阳集热器的热性能参数受到集热器透明盖板层数以及种类的影响，且将集热器瞬时效率和排管内工质出口温度视为目标函数，进而研究了集热面积、排管内径、排管间距和排管内总工质流量的优化问题。

An experimental and numerical investigation of

Combustion and Flame 145(2006)740–764/locate/combustﬂameAn experimental and numerical investigation of n -heptane/air counterﬂow partially premixed ﬂamesand emission of NO x and PAH speciesPaolo Berta a ,Suresh K.Aggarwal a ,∗,Ishwar K.Puri ba Department of Mechanical and Industrial Engineering,University of Illinois at Chicago,Chicago,IL,USAb Department of Engineering Science and Mechanics,Virginia Polytechnic Institute and State University,Blacksburg,VA,USAReceived 13July 2005;received in revised form 27January 2006;accepted 30January 2006Available online 23March 2006AbstractAn experimental and numerical investigation of counterﬂow prevaporized partially premixed n -heptane ﬂames is reported.The major objective is to provide well-resolved experimental data regarding the detailed structure and emission characteristics of these ﬂames,including proﬁles of C 1–C 6,and aromatic species (benzene and toluene)that play an important role in soot formation.n -Heptane is considered a surrogate for liquid hydrocarbon fuels used in many propulsion and power generation systems.A counterﬂow geometry is employed,since it provides a nearly one-dimensional ﬂat ﬂame that facilitates both detailed measurements and simulations using comprehen-sive chemistry and transport models.The measurements are compared with predictions using a detailed n -heptane oxidation mechanism that includes the chemistry of NO x and PAH formation.The reaction mechanism was syn-ergistically improved using pathway analysis and measured benzene proﬁles and then used to characterize the effects of partial premixing and strain rate on the ﬂame structure and the production of NO x and soot precursors.Measurements and predictions exhibit excellent agreement for temperature and major species proﬁles (N 2,O 2,n -C 7H 16,CO 2,CO,H 2),and reasonably good agreement for intermediate (CH 4,C 2H 4,C 2H 2,C 3H x )and higher hydrocarbon species (C 4H 8,C 4H 6,C 4H 4,C 4H 2,C 5H 10,C 6H 12)and aromatic species (toluene and benzene).Both the measurements and predictions also indicate the existence of two partially premixed regimes;a double ﬂame regime for φ<5.0,characterized by spatially separated rich premixed and nonpremixed reaction zones,and a merged ﬂame regime for φ>5.0.The NO x and soot precursor emissions exhibit strong dependence on partial premixing and strain rate in the ﬁrst regime and relatively weak dependence in the second regime.At higher levels of partial premixing,NO x emission is increased due to increased residence time and higher peak temperature.In contrast,the emissions of acetylene and PAH species are reduced by partial premixing because their peak locations move away from the stagnation plane,resulting in lower residence time,and the increased amount of oxygen in the system drives the reactions to the oxidation pathways.The effects of partial premixing and strain rate on the production of PAH species become progressively stronger as the number of aromatic rings increases.©2006The Combustion Institute.Published by Elsevier Inc.All rights reserved.Keywords:n -Heptane;Partially premixed ﬂames;NO x and PAH species measurements;Detailed modeling*Corresponding author.E-mail address:ska@ (S.K.Aggarwal).P.Berta et al./Combustion and Flame145(2006)740–7647411.IntroductionA major portion of the world’s energy demands is currently met by the combustion of liquid fuels. Closely associated with the beneﬁts derived from combustion are the hazards it causes to human life and environment.The products of combustion of most commercially available fuels contain pollutants such as particulate matter,unburned and partially unburned hydrocarbons,carbon monoxide,and oxides of nitro-gen and sulfur.These pollutants have many harmful effects including speciﬁc health hazards,acid rain, smog,global warming,and ozone depletion.The ac-ceptability of a new grade of fuel or design of a new combustion system at present depends as much on its emission characteristics as on its combustion ef-ﬁciency.Consequently,energy conservation and en-vironmental concerns provide a strong motivation for fundamental studies on the mechanism of soot and NO x formation inﬂames.Partially premixedﬂames contain a rich premixed fuel–air mixture in a pocket or stream,and,for com-plete combustion to occur,they require the transport of oxidizer from an appropriately oxidizer-rich(or fuel-lean)mixture that is present in another pocket or stream.Partial oxidation reactions occur in fuel-rich portions of the mixture and any remaining unburned fuel and/or intermediate species are consumed in the oxidizer-rich portions.Partially premixedﬂames are important in numerous applications.They are rele-vant to turbulent nonpremixed combustion,which can contain regions where local extinction occurs,fol-lowed by partial premixing and reignition.Partially premixed combustion plays a fundamental role in the stabilization of lifted nonpremixedﬂames in which propagating premixed reaction zones anchor a non-premixed reaction zone.In addition,in most liquid-fueled combustion devices,such as internal combus-tion engines,industrial furnaces,and power station gas turbines,the fuel is introduced in the form of a spray of fuel droplets of different sizes.The smaller droplets evaporate at a much higher rate than the larger ones.The resulting fuel vapor mixes with air, forming locally fuel-rich zones.The larger droplets then burn in this mixture in a partially premixed mode.Partially premixedﬂames may also result in lean direct injection diesel engines.The liquid fuels that are used in internal com-bustion engines and gas turbines are typically blends of several components.Generally,fuels with desired properties are prepared by mixing expensive volatile components with cheaper heavier fuels.The detailed simulation and analysis ofﬂames burning these fu-els in actual engines is a prohibitively complex task single-component or bicomponent fuel,based on the most abundant species in the actual fuel.In prac-tical liquid fuels such as gasoline and diesel fuels, n-C7H16is relatively abundant,and hence often used as a surrogate for these fuels.Since the soot-and NO x-forming mechanisms are closely related to the chemical kinetics and structure ofﬂames,a detailed study of partially premixed n-heptaneﬂames(PPFs)is of direct relevance to op-timizing the operating conditions of a diesel engine for minimum production of soot,unburned hydro-carbons,and NO x.Due to these diverse applications and fundamental relevance,partially premixedﬂames have been investigated extensively in recent years. However,the bulk of these studies have focused on methane–airﬂames[1–5],motivated perhaps by the fact that detailed reaction mechanisms are available to model the methane–air chemistry.With the excep-tion of some recent investigations[6–9],the literature regarding the burning of higher hydrocarbon fuels,es-pecially liquid fuels,in partially premixedﬂames is relatively sparse.Li and Williams[6]reported mea-surements of several major and intermediate species in n-heptane PPFs burning a droplet/air fuel mixture in a counterﬂow conﬁguration.Seiser et al.[10]re-ported an experimental investigation of prevaporized n-heptane counterﬂow nonpremixedﬂames.Xue and Aggarwal[7]characterized the structure of n-heptane counterﬂow PPFs through a numerical investigation and subsequently investigated the effect of double ﬂame structure on NO x formation in theseﬂames[8]. Berta et al.[38]recently reported an experimental and numerical investigation of the structure and emis-sion characteristics of prevaporized n-heptane non-premixedﬂames in a counterﬂow conﬁguration.Our literature indicates that there is a lack of de-tailed experimental data pertaining to the structure and emission characteristics of n-heptane PPFs.This is rather surprising since n-heptane has been con-sidered a good surrogate for liquid fuels used in many practical combustion systems,and its oxidation chemistry has been extensively investigated.More-over,compared to other combustion systems,includ-ing premixed and nonpremixedﬂames,a PPF pro-vides a more stringent crucible for the validation of reaction mechanisms[11].This is due to the exis-tence of multiple reaction zones and interactions be-tween them involving both chemistry and transport processes.These interactions also play a signiﬁcant role in determining the NO x and soot emissions from theseﬂames.Motivated by the above considerations,we report herein an experimental–computational investigation of partially premixed n-heptaneﬂames established742P.Berta et al./Combustion and Flame145(2006)740–764species concentrations,especially those of C1–C6hy-drocarbons,for a wide range of partial premixing (i.e.,equivalence ratios)and strain rates.C1–C6hy-drocarbons are key intermediates in the fuel decom-position pathway and their characterization is crucial for understanding the combustion of heavier fuels,es-pecially in the context of partially premixedﬂames, which are hybridﬂames and whose structure is char-acterized by both transport and chemical kinetics. Species concentration proﬁles of intermediate hydro-carbons can be subsequently used for the validation of computational models and reaction mechanisms involving simulations of liquid fuels in general,and n-heptane in particular.Therefore,we report well-resolved measurements of major species(n-C7H16, O2,N2,CO2,and H2O),intermediate species(CO, H2,CH4,C2H4,C2H2,and C3H x),higher hydrocar-bon species(C4H8,C5H10,and C6H12),and the ma-jor soot precursor(benzene)over a large parametric space characterized in terms of equivalence ratio(φ) and strain rate(a G).The measurements also focus on the resolution of unsaturated C3and C4species such as propene,propyne,allene,butene,1,3-butadiene, 1-buten-3-yne,1,3-butadiyne,and aromatic species (benzene and toluene).Some of these species have never been previously measured for n-heptane coun-terﬂow PPFs.Another objective is to characterize the effect of partial premixing on the formation of NO x and soot precursors,such as acetylene,benzene,and other PAH(polycyclic aromatic hydrocarbon)species,in n-heptane PPFs.Acetylene represents a key species in the formation of polyaromatic structures through the hydrogen abstraction carbon addition(HACA)mech-anism[12],while benzene represents the simplestaromatic molecule.The numerical investigation hasbeen performed using a detailed mechanism that is ca-pable of simulating the formation of NO x and PAHsup to coronene.2.The experimental setupA schematic of the experimental setup used to es-tablish prevaporized n-heptane counterﬂowﬂames ispresented in Fig.1.A mixture of prevaporized n-heptane and nitrogen fuel was introduced from thebottom nozzle.A nitrogen curtain was establishedthrough an annular duct surrounding the fuel jet inorder to isolate theﬂames from ambient disturbances.This nitrogen and combustion products were ventedand cooled through another annular duct around theoxidizer nozzle.The diameter of each nozzle was27.38mm,and the separation distance(L)betweenthem was varied from10to20mm.The veloci-ties of the two streams deﬁne the global strain rateas a G=(2|V O|/L)(1+(|V F|/|V O|)(ρF/ρO)1/2)[13] and were chosen to satisfy the momentum balance,ρO V2O=ρF V2F.Hereρrepresents density,V gas ve-locity,and the subscripts O and F refer to oxidizer andfuel nozzles,respectively.The oxidizer was air at room temperature,whilethe fuel stream consisted of mixtures of air and pre-vaporized n-heptane.The fuel nozzle was heatedand its temperature controlled to maintain the fuel-containing stream at a400K temperature at theburnerP.Berta et al./Combustion and Flame145(2006)740–764743exit.In the bottom part of the burner preheated air was mixed with the pure fuel stream to form a fuel–air mixture of the desired equivalence ratio.The n-heptane vapor was formed in a prevaporizer,which was an electrically heated stainless steel chamber.The desired massﬂow rate of n-heptane into the prevapor-izer was maintained by a liquid pump.Approximately three-fourths of the chamber wasﬁlled with glass beads to impede theﬂow,thereby increasing its res-idence time and thus enhancing the heat transfer to the liquid fuel.The temperature of the fuel vapor ex-iting the chamber was monitored by a thermocouple.Temperature proﬁles of variousﬂames were mea-sured using a Pt–Pt13%Rh thermocouple with a spherical bead diameter of0.25mm and wire diam-eter of0.127mm.The measured values were cor-rected for radiation heat losses from the bead,assum-ing a constant emissivity of0.2and a Nusselt num-ber of2.0[10].Species concentration proﬁles were measured using a Varian CP-3800gas chromatograph (GC).Samples were drawn from theﬂame with a quartz microprobe that had a0.34-mm tip diameter and0.25-mm tip oriﬁce.Constant vacuum was ap-plied at the end of the line through a vacuum pump. The line carrying the sample to the GC was made of fused silica and was heated to prevent conden-sation.A portion of the sample was injected into a Hayesep DB100/120packed column connected to a thermal conductivity detector to measure light gases (up to C2H4)and another into a Petrocol DH capillary column that was placed inline with aﬂame ioniza-tion detector to obtain hydrocarbon distributions up to C7H16.The temperature in the gas chromatograph oven was gradually increased to minimize the analy-sis time.The temperature and pressure in the sam-pling loops were controlled to ensure that the same volume of gas was sampled for each analysis.The chromatogram peaks have been converted into mole fractions with calibration constants that were obtained separately for every species from known standards. Water molar fractions were obtained through a mass balance of carbon and hydrogen atoms.The errors in measurement of the liquid fuel and airﬂow rates are within5%,leading to an uncertainty of about5% in equivalence ratio.The compositions of both the fuel and air streams were also measured using GC. C3and C4unsaturated species were measured ofﬂine by an HP6890gas chromatograph connected to a mass spectroscopy detector.The sample was collected in a stainless steel vessel.The whole line and vessel were heated to minimize condensation.Temperature programming was employed to reduce the analysis time.The temperature and pressure in the sampling loop were controlled and measured to ensure that the fractions with calibration constants that were obtained separately for every species from known standards. The uncertainties in GC measurements are between 5%and10%depending on the species.3.The physical–numerical modelMost of the studies on heptaneﬂames reported in the literature deal with nonpremixedﬂames.Experi-mental results have been obtained in several conﬁg-urations:liquid pool burners[14,15],droplet burn-ing[16,17],and premixedﬂames[18].n-Heptane combustion chemistry has been investigated on many different levels.One-step global and reduced mech-anisms[19,20]have been empirically derived toﬁt experimental data of burning velocities andﬂame extinction.Held et al.[17]reported a semidetailed mechanism and validated it usingﬂow reactor,shock tube,stirred reactor,and laminarﬂame speed experi-mental data.The mechanism was subsequently used for predicting ignition delays in shock tubes[21] and for numerical investigations of partially premixed ﬂames[7,8].Lindstedt and Maurice[22]developed a detailed n-heptane mechanism,addressing in detail the H abstraction reactions on the C7molecule and its decomposition into smaller fragments.The mecha-nism was improved in subsequent work[23]to further characterize the formation and oxidation of aromatic molecules.Detailed n-heptane mechanisms have also been reported by Chakir et al.[24],Curran et al.[25], and Babushok and Tsang[26].The kinetic mechanism(SOX)used to model n-heptaneﬂames in the present study was previously developed by extending a detailed oxidation scheme for several fuels[27,28].Due to the hierarchical mod-ularity of the mechanistic scheme,this model is based on a detailed submechanism of C1–C4species.As-suming analogy rules for similar reactions,only a few fundamental kinetic parameters are required for the progressive extension of the scheme toward heavier species.The resulting kinetic model of hydrocarbon oxidation from methane up to n-octane consists of about170species and5000reactions.We have selected this mechanism for our simu-lations since the subset of n-heptane oxidation reac-tions included in it has been extensively tuned using experimental measurements for pure pyrolysis condi-tions,oxidation in jet-stirred and plug-ﬂow reactors, and shock-tube experiments[29].Moreover,a rela-tively detailed model for polycyclic aromatic hydro-carbons(PAHs)that are soot precursors is contained in the mechanism.The formation of theﬁrst aromatic rings by C2and C4chemistry and by resonantly sta-744P.Berta et al./Combustion and Flame145(2006)740–764 gated[28,30].Further growth of PAH species up tocoronene(C24H12)is also modeled through the well-known HACA mechanism[31],which has been ex-tensively validated for counterﬂowﬂames burning avariety of fuels[32].The main consumption reactionsof aromatics and PAHs are H abstraction reactions byH and OH radicals.The high-temperature reactionshave been validated against substantial experimentaldata[27,28,30].Numerical simulations of counterﬂowﬂames wereperformed using the OPPDIF code[33],which is ca-pable of modeling combustion between two opposedjets.The code was modiﬁed to handle the complexreaction mechanism and to account for thermal radia-tion through an optically thin model[34].Most ther-modynamic properties were obtained from Burcat andMcBride[35],and unavailable properties were esti-mated using the group additivity and difference meth-ods[36].Transport properties were obtained from theCHEMKIN database[37]wherever available,whileunavailable data were deduced through analogy withknown species.To establish grid independence,numerical solu-tions were obtained on increasinglyﬁner grids,andby changing GRAD and CURV parameters,until novariation was observed between two grid systems.4.Results and discussionTo perform a detailed experimental and numericalinvestigation of theﬂame structure and emission char-acteristics,prevaporized n-heptane PPFs were estab-lished at different strain rates(a G)and equivalence ra-tios(φ).Table1shows the parametric space in termsof a G andφfor seven PPFs,designated as FlamesA–G,which are analyzed experimentally and numer-ically in the present study.For all the cases,the fuelstream was introduced from the bottom nozzle and theoxidizer from the top nozzle.The oxidizer was pureair,while the fuel stream was a mixture of n-heptaneand air with the desired value ofφ.Note that PPFsTable1Operating conditions in terms of strain rate,equivalence ra-tio,and nozzle separation distance for the cases investigatednumerically and experimentallyFlame Strain rate(s−1)EquivalenceratioNozzle separation(cm)A5015.31 B506.11 C502.52 D1008.01 E15012.61established at a G=100s−1and different values ofφhave been investigated in our previous work[9].Con-sequently,only one value ofφis considered at this strain rate(Flame D).For preliminary analysis,digital images of several PPFs were taken for different values of strain rate, partial premixing,and nozzle separation distance.The images of four representativesﬂames,i.e.,Flames A, C,G,and E,are presented in Fig.2.The images of Flames A,G,and E were taken at the same exposure time,while that of Flame C was taken at double the exposure time for it was less luminous.For Flame A, which is characterized by low strain rate and low level of partial premixing,an orange-red zone can be ob-served below the familiar green-blue doubleﬂame structure,with green from the C2chemiluminescence in the premixed zone and blue from the CO oxidation in the nonpremixed zone.Even though the equiva-lence ratio is high theﬂame does not appear as sooty as a nonpremixedﬂame,which is bright yellow.The red zone disappears as the strain rate and/or level of partial premixing are increased.Flame C shows the greatest separation between the two reaction zones; as the partial premixing approaches stoichiometric conditions the premixedﬂame moves closer to the fuel nozzle.Thisﬂame does not appear asﬂat as the othersﬂames because the nozzle separation dis-tance had to be increased to obtain the desired strain rate.The doubleﬂame structure can still be seen in Flame G,which appears brighter,since more fuel is consumed during the same exposure time due to the higher strain rate.In Flame E,which is characterized by a higher equivalence ratio,the doubleﬂame struc-ture can barely be noticed,as the two reaction zones are nearly merged.A detailed comparison of measurements and sim-ulations for the sevenﬂames listed in Table1is pre-sented in Figs.3–9.Eachﬁgure shows the temper-ature,axial velocity,and species mole fraction pro-ﬁles.The predictions are shown by continuous lines, while the experimental data are shown by symbols. The three vertical lines in eachﬁgure indicate that(1) the nonpremixed reaction zone location that is iden-tiﬁed by the peak in temperature proﬁle and marked by the solid vertical line;(2)the stagnation plane that is marked by the dashed vertical line;and(3)the rich premixed zone location that is identiﬁed by the peak in hydrogen proﬁle and marked by the vertical dotted line.The experimental proﬁles are highly re-solved,since particular effort has been expended to capture the regions characterized by high chemical activity and steep gradients.Moreover,measurements of several intermediate hydrocarbon species that areP.Berta et al./Combustion and Flame145(2006)740–764745Fig.2.Digital images of partially premixed n-heptaneﬂames(Flames A,C,G,E).(1)A general observation from the measured andpredicted proﬁles for the sevenﬂames is that PPFs are characterized by a doubleﬂame struc-ture:a rich premixed zone is established down-stream of the fuel nozzle and characterized by pyrolysis and partial oxidation of n-heptane.The products of partial oxidation,namely CO,H2, and intermediate hydrocarbon species,are trans-ported and consumed in the nonpremixed reac-tion zone located on the oxidizer side.The double ﬂame structure becomes visually more distinct as a G decreases and/or the level of partial pre-mixing increases(i.e.,φdecreases).This also increases the separation distance between the two reaction zones.The premixed reaction zonevelocity(V x)matches the burning velocity(S L)of the stretchedﬂame.Since S L increases asφisreduced,the premixedﬂame moves away fromthe stagnation plane toward the fuel nozzle tosatisfy the condition S L=V x.The nonpremixed ﬂame is established on the oxidizer side at the lo-cation(x n)where the intermediate fuel speciesand oxidizerﬂuxes are transported in stoichio-metric proportion.Therefore,the separation dis-tance between the two reaction zones increasesas the level of partial premixing is increased.In-creasing the strain rate has the opposite effect,since for largerﬂow velocities the location x p ispushed toward the stagnation plane.(2)For all the sevenﬂames analyzed,there is gen-746P.Berta et al./Combustion and Flame145(2006)740–764Fig.3.Predicted(lines)and measured(symbols)proﬁles for Flame A.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles ofP.Berta et al./Combustion and Flame145(2006)740–764747Fig.4.Predicted(lines)and measured(symbols)proﬁles for Flame B.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles of benzene.The vertical lines in some of theﬁgures indicate the locations of the stagnation plane,the premixedﬂame,and the748P.Berta et al./Combustion and Flame145(2006)740–764Fig.5.Predicted(lines)and measured(symbols)proﬁles for Flame C.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles of benzene.The vertical lines in some of theﬁgures indicate the locations of the stagnation plane,the premixedﬂame,and theP.Berta et al./Combustion and Flame145(2006)740–764749Fig.6.Predicted(lines)and measured(symbols)proﬁles for Flame D.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene, acetylene,and C3hydrocarbons;mole fraction proﬁles of C4H8,C5H10,and C6H12oleﬁns;and mole fraction proﬁles of benzene.The vertical lines in some of theﬁgures indicate the locations of the stagnation plane,the premixedﬂame,and theFig.7.Predicted(lines)and measured(symbols)proﬁles for Flame E.Temperature and axial velocity proﬁles;mole fraction proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene,acetylene,and C hydrocarbons;mole fraction proﬁles of C H,C H,and C H oleﬁns;and mole fraction proﬁles ofproﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene,proﬁles of O2,N2,and n-C7H16;mole fraction proﬁles of H2O,CO2,CO,and H2;mole fraction proﬁles of CH4,ethylene,dictions.The predictions reproduce the measured partially premixedﬂame structure for all the strain rates and equivalence ratios investigated.Both measurements and predictions indicate that at low level of partial premixing(highφ)and/or high strain rate,the two reaction zones are nearly merged,and that as the level of partial premixing increases and/or a G decreases,the separation dis-tance between the two reaction zones increases and the doubleﬂame structure becomes more dis-cernible.(3)There is good quantitative agreement betweenmeasurements and predictions for major reac-tant and product species(n-heptane,O2,N2,and CO2)as well as for intermediate fuel species(H2 and CO).A good agreement between the mea-sured and predicted peak concentrations of these species and between the locations of their peak concentrations implies that both the transport and chemistry are reasonably well reproduced in the simulations.For instance,a good agreement be-tween the measured and predicted locations of the H2and CO concentration peaks implies that the location of the rich premixed zone is well pro-duced by the simulations.Similarly,good agree-ment between the measured and predicted loca-tions of the CO2concentration peaks indicates that the location of the nonpremixed zone is well reproduced by the simulations.The measured and predicted temperature proﬁles also exhibit good agreement,although there is a mismatch between the locations of the respective peaks.A similar discrepancy has been observed by otherinvestigators,and may partly be attributed to the catalytic effect of the thermocouple.(4)The n-heptane and O2proﬁles on the fuel side in-dicate that the reaction mechanism underpredicts the consumption rates of these species in the rich premixed zone;it can be seen forﬂame E and toa lesser extent also for Flames A,B,F,and G.This is further corroborated by the H2,CO,and intermediate hydrocarbon species(C2H4,C2H2, C4H8,and C5H10)proﬁles.This discrepancy be-comes less pronounced,however,as the level of partial premixing is increased,i.e.,asφis de-creased.(5)There is also fairly good quantitative agreementbetween measurements and predictions for light hydrocarbon species(CH4,C2H4,C2H2).How-ever,the quantitative agreement deteriorates, although there is good qualitative agreement, for higher hydrocarbon species(C4H8,C5H10), which are present at relatively low concentra-the measured proﬁles of this species are shown for all the sevenﬂames.(6)The comparison of the measured and predictedH2O proﬁles exhibits large discrepancies.This is due to the fact that H2O concentration is not measured directly from GC;it is obtained by ap-plying a balance of carbon and hydrogen atoms using the GC measurements of the other species.This procedure implies equal diffusivity for all the species,an assumption that is not well satis-ﬁed,especially on the fuel side due to the pres-ence of large(C7H16)and small(H2)molecules that have very different diffusion coefﬁcients. (7)In the present study,particular attention wasgiven to the measurement and prediction of ben-zene proﬁles,since this species has the simplest structure,with a single aromatic ring,and rep-resents perhaps the most important intermediate in the growth process to PAHs and soot.In spite of their relatively low concentrations,the pre-dicted and measured benzene proﬁles exhibit fairly good agreement for all the sevenﬂames shown in Figs.3–9.Both measurements and pre-dictions indicate that the benzene concentration decreases as the level of partial premixing and/or strain rate is increased.It is worth mentioning that the predictions are based on the reaction mechanism that was synergistically improved using pathway analysis and measured benzene proﬁles in our previous investigation[38].In order to better characterize the pyrolysis zone, further analysis was performed for Flames A,E,and G and the results are presented in Figs.10–12.The objective was to obtain quantitative data on the dis-tribution of unsaturated C3–C4intermediates for dif-ferent levels of partial premixing(equivalence ratio) and strain rates.The C3–C4intermediates proﬁles presented here are extremely valuable for the develop-ment and testing of n-heptane reaction mechanisms, since these species constitute the main decomposi-tion products of the C7H15radical.In addition,they directly affect the formation of the propargyl radical [26],which,through benzene and PAHs,leads to soot formation.The measurements were taken using the ofﬂine technique,as described earlier,on the fuel side of theﬂame,where n-heptane undergoes rapid con-sumption.For most of the species reported in Figs. 10–12,there is good agreement between predictions and measurements,especially considering that the de-tailed chemistry model has not been tuned for this set of data,implying that the reaction mechanism ade-。

基于PLC西门子S7-200的带式输送机控制系统设计

基于PLC西门子S7-200的带式输送机控制系统设计李博【摘要】Belt conveyor is a conveyor belt driven by the driving roller, as a continuous conveying equipment, which is one of the main transport equipment in modern mines. In particular, the modern large-scale coal mines, coal transportation is mainly through the conveyor to complete. Conveyor is characterized by large transmission capacity, small power consumption, simple structure, strong adaptability to the material, easy to form water production line, so that the enterprise production process to achieve mechanization. Belt conveyor in coal mine and coal preparation plant plays a important role, under the background of enterprise development of coal mining mechanization, automation, automatic control operation of the conveyor belt is particularly important. In this paper, the design and programming of the conveyor system are carried out with the SIEMENS S7-200 programmable controller as an example.%带式输送机是一种由驱动滚筒带动的输送带,作为一种连续输送的设备,其是现代矿井的主要运输设备之一.尤其是现代化大型的煤矿,煤的运输主要通过输送机来完成.输送机的特点是输送能力大,功耗小,构造简单,对物料的适应性强,便于组成流水生产线,使企业生产过程实现机械化.带式输送机在煤矿及洗煤厂发挥着重要作用,在企业发展采煤机械化、自动化的背景下,带式输送机的自动控制运行显得尤为重要.本文以西门子S7-200可编控制器为例进行输送机系统的编程和设计.【期刊名称】《现代制造技术与装备》【年(卷),期】2015(000)006【总页数】2页(P21-22)【关键词】带式输送机;S7-200;可编控制器【作者】李博【作者单位】河南龙宇能源机电制修厂,商丘 476600【正文语种】中文以西门子S7-200可编控制器为控制系统的核心，结合若干监测传感器的配合应用，最终构成带式输送机的控制系统。

Experimental and Numerical Studies of Failure Mode

Transactions of Tianjin University (2018) 24:387–400https:///10.1007/s12209-018-0123-0E xperimental and Numerical Studies of Failure Modesand Load-Carrying Capacity of Through-Diaphragm ConnectionsB in R ong1,2 · S huai L iu 1 · Z henyu L i 1 · R ui L iu 1R eceived: 15 March 2017 / Revised: 20 May 2017 / Accepted: 21 August 2017 / Published online: 3 February 2018© T ianjin University and Springer-Verlag GmbH Germany, part of Springer Nature 2018A bstractS hear failure in panel zones and plastic hinges in steel beams are the two major failure modes of connections between concrete-ﬁ lled steel tubular (CFST) columns and steel beams. To investigate the behavior of this type of connection in both modes, two through-diaphragm connections were tested under cyclic and monotonic loadings and the load-carrying capac-ity, ductility, and strength of degradation of connections were discussed. Using ABAQUS software, we developed nonlinear ﬁ nite-element models (FEMs) to simulate the load-carrying capacity and failure modes of the connections under monotonic loading. The ﬁ nite-element (FE) analysis and test results showed reasonable agreement for the through-diaphragm con-nections, which conﬁ rms the accuracy of FEMs in predicting the load-carrying capacity and failure modes of connections. Based on the validated FEM, a parametric study was then conducted to investigate the inﬂ uence of the thicknesses of the tube and diaphragm on the load-carrying capacity and failure modes of these connections. The results indicate that the strength, stiﬀ ness, and load-carrying capacity are inﬂ uenced less by the thickness of the diaphragm, and more by the thickness of the steel tube. According to the FE analysis results, it can be found that the critical condition between the two failure modes is determined by the shear resistance and bending resistance.K eywords T hrough-diaphragm connections · F ailure mode · F inite-element analysis · P arametric analysis · M onotonic loading · C yclic loadingI ntroductionC oncrete-ﬁ lled steel tubular (CFST) columns have been extensively utilized in high-rise buildings owing to their practical advantages. The tube in a CFST column serves as a convenient formwork for concrete and provides external conﬁ nement for the cured concrete. By conﬁ ning concrete in a CFST column, an increase in the concrete’s compressive strength as well as preventing the concrete from spalling can be realized when subjected to overloading. Furthermore, the concrete inside the tube serves to restrain the occurrence of local buckling in the web of the steel tube. The connec-tions between CFST columns and steel beams are critical elements, and their seismic behavior plays an important role in the security of structural design.A lot of experimental and analytical studies have been conducted with respect to diﬀ erent types of connections between CFST columns and steel H-shaped beams. Based on cyclic tests of ten through-diaphragm connections, Morino et al. [ 1] reported two failure modes for these three-dimensional connections, i.e., panel zone and column bend-ing failures. Sasaki et al. [ 2] developed an analytical model for the shear resistance of these connections based on the virtual work principle. Kang et al. [ 3] tested eight connec-tions with T-stiﬀ eners and reported their stable hysteresis behavior in beam failure mode. Cheng and Chung [ 4] tested ﬁ ve circular CFST connections to validate a proposed non-linear force–deformation model, but all the specimens failed in welding fractures. Nishiyama et al. [ 5] performed cyclic tests for through-diaphragm and exterior-diaphragm connec-tions using high-strength steel and concrete and observed shear failure in the test process; they also developed an ana-lytical model for the shear resistance of connections. Nie et al. [ 6] tested 14 connections between CFST columns and*Z henyu L il izhenyu@1S chool of Civil Engineering,T ianjin University,T ianjin 300072,C hina2 K ey Laboratory of Coast Civil Structure Safety ,M inistryof Education ,T ianjin 300072 ,C hina1 3388 B. Rong et al.1 3steel–concrete composite beams, and reported that exterior-diaphragm connections had better seismic performance. Qin et al. [ 7 , 8 ] compared the seismic behavior of an improved type of through-diaphragm connections, which featured a tapered ﬂ ange under cyclic loading, and found welding failure and beam failure to be their main failure modes. In the above studies, beam failure, welding failure, and shear failure of the panel zone comprised the three main failure modes for all connection types. Although some analyti-cal models have been proposed regarding the resistance of connections, the critical condition for these failure modes remains uncertain. W ith the rapid progress in computer technology, ﬁ nite-element analysis (FEA) has recently developed into an alter-native approach for investigating the mechanical properties of connections between CFST columns and steel beams. Chiew et al. [ 9 ] used Marc software to establish a ﬁ nite-ele-ment model (FEM) to study the key parameters in moment resistance of connections between concrete-ﬁ lled circular steel tubular columns and I-shaped steel beams. Kang et al. [ 3 ] used ABAQUS software to conduct a nonlinear FEA of the hysteretic behavior of CFST-column-to-H-beam connec-tions with T-stiﬀ eners and penetrated elements. Shin et al. [ 10 ] improved this simulation by considering cyclic loading and eventually found a good correlation between predicted and measured load–displacement hysteretic curves. Choi et al. [ 11 ] studied the moment–rotation relationship between square CFST columns and steel-beam joints using ANSYS. Chou et al. [ 12 ] used ABAQUS to perform an analytical study on the seismic performance at moment-resisting joints between post-tensioned steel beams and CFST columns. Zhang et al. [ 13 ] used ANSYS to conduct parametric stud-ies on the inﬂ uence of axial load ratio on the shear behavior of through-diaphragm connections, indicating that the axial load ratio should be limited to < 0.4 and its inﬂ uence should be considered in connection analysis and design. Nie et al. [ 14 ] employed ANSYS to perform a correlation study and also conducted parametric studies to investigate the eﬀ ects of axial load level, width-to-thickness ratio, and dimensions of exterior diaphragms on the load–displacement response. Zhang et al. [ 15 ] used ABAQUS to investigate the seismic performance of through-diaphragm connections between rectangular CFST columns and steel beams.M ost of the numerical researches were limited to a single failure mode, and a few researchers discussed the criticalcondition associated with the shear failure in the panel zone and the bending failure in the steel beam. The key param-eters that determine these failure modes have not been found yet. In this paper, we presented experimental and numerical studies on the failure mode of through-diaphragm connec-tions between rectangular concrete-ﬁ lled steel tubular col-umns and steel beams. Based on the results of parametric studies, a critical condition between the two failure modes was also summarized.E xperimental StudyD esign of Specimens Two specimens were designed to investigate the failure modes of T-shaped through-diaphragm connections. Speci-men T1 was designed with a smaller section size to evalu-ate the shear capacity of the panel zone. Specimen T2 was prepared for bending failure in beam. Cyclic loading and monotonic loading were applied to T1 and T2, respectively. Table 1 lists the dimensions of each specimen and Fig. 1 shows the details of the panel zone.T he steel tubes of specimens were square, and the steel beams were H-shaped. The steel tubes were discontinuous at the positions of diaphragms, and the diaphragms cut through the steel tube columns. All the steel components of speci-mens were connected by full-penetration butt welds with backing bars that had been processed and welded together in one factory in advance. The strength grade of all steel was Q235B. Then, the tube column was ﬁ lled with C40 concrete. Standard tensile pieces of steel and standard test cubes of concrete were prepared for material property test. Tables 2 and 3 list the material properties of steel and con-crete, respectively.T est Setup F igure 2 shows a schematic drawing and photo of the test setup. The T-shaped specimens were placed horizontally on the testing frame. The ends of columns were connected to the ﬁ xed hinges. The right hinged support was designed as a long hole, so that it would only provide vertical force for the purpose of the simulation of simply supported constraint.T able 1D imension ofspecimensS pecimen C olumn section (mm) B eam section (mm) C olumn section in panel zone (mm) T hickness ofdiaphragm(mm)G rade ofconcrete T 1 ᦕ200 × 200 × 12 H 250 × 200 × 8 × 12 ᦕ200 × 200 × 6 14 C 40 T 2 ᦕ250 × 250 × 12 H 250 × 250 × 8 × 14 ᦕ250 × 250 × 12 14 C 40389Experimental and Numerical Studies of Failure Modes and Load-Carrying Capacity of… 1 3D uring the actual experiment, the horizontal load at thebeam end was controlled by the force–displacement hybrid control loading system. The force-control was adopted before the yielding of the specimen and displacement-control was adopted after yielding. Stepwise force load was applied to the beam end and repeated one time at each step loading before yielding. After yielding, the loading on T1 was controlled by multiplying the horizontal displacementof beam end, and it was repeated three times for each step. The loading on T2 was monotonic and under displacement-control. To make the deformation of the specimens under loading reach basic stability, experimental data were col-lected after keeping the loading for 5 min per level. The con-trolled displacement in the loading process was measured by linear variable diﬀ erential transformer (LVDT), which was placed at the end of the beam corresponding to the location of horizontal actuator.F ailure Phenomena Both specimens behaved in a ductile manner. The cyclic and monotonic loading proceeded smoothly under control. F or specimen T1, the thickness of tube in the panel zone decreased to 6 mm. Before yielding, the load–displacement curve was linear. At a load of 120 kN, the panel zone yielded and no signiﬁ cant phenomenon was observed. Beyond this point, the slope of the load–displacement curve became smaller. Shear deformation started in the webs of the panel zone upon loading into the second yielding displacement cycle. With the increasing displacement of the beam end, the horizontal load continued to increase, and the tube web began to bear more shear stress transferred from the dia-phragms. Then, local buckling occurred on the tube webF ig. 1 D etails of panel zone (unit: mm). a T1 and b T2T able 2M aterial properties of steel T hickness of steel (mm) Y ield strength f y (N/mm 2) T ensilestrength f u (N/mm 2 ) 6268.9 387.9 8249.6 373.1 12 313.8 439.8 14 252.2 414.6 T able 3M aterial properties of concrete G rade of concrete f cu (N/mm 2) f ′c (N/mm 2 ) C 40 43.55 34.40390 B. Rong et al.1 3and the shear deformation became obvious, as shown in Fig. 3 a . The load reached its peak value of 198.8 kN, cor-responding to a 52.5-mm displacement. As the displacement increased, the load dropped to 85% of its peak value, and the tube web in the panel zone started to fail. The specimen was then unloaded and the test was terminated. It can be seen that the column web of panel zone ﬁ nally bulged outward in the test. Shear failure mode could be identiﬁ ed in T1, with an obvious local buckling in the tube web, as shown in Fig. 3 b . In contrast, the steel beam experienced a little bending deformation.F or specimen T2, the thicknesses of all the tubes were 12 mm. Monotonic loading proceeded smoothly under con-trol. At a load of 211.5 kN, the beam yielded and the slope of load–displacement curve became smaller. The loading then entered the displacement-control stage. T2 reached its ultimate capacity of 275.1 kN with a corresponding dis-placement value of 76.2 mm. At four times of the yielding displacement, the left beam ﬂ ange buckled in compression and the right one lengthened in tension, after which the load-carrying capacity decreased slowly. At ﬁ ve times of the yielding displacement, a crack occurred at the weld line between the beam web and steel tube, as shown in Fig. 4 a .The test ended when the load dropped to 85% of its ultimate value. Little shear deformation was observed in the panel zone during all the loading history. Bending failure mode with a plastic hinge in the beam ﬂ ange could be identiﬁ ed in T2,as shown in Fig. 4b .F ig. 2 T est setup. a Schematic drawing of test setup (unit: mm); b photo of test setupF ig. 3 E xperimental phenom-ena of T1. a Shear deformation of panel zone; b local buckling of panel zoneF ig. 4 E xperimental phenomena of T2. a Weld failure on beam web;b buckling on beam ﬂ ange391Experimental and Numerical Studies of Failure Modes and Load-Carrying Capacity of… 1 3D iscussion of Test Results L oad–Displacement Relationship F igure 5 shows the load–displacement hysteretic loops of T1. These loops are smooth and full, indicating good duc-tile behavior with high energy absorption by specimen T1. After the load reached its maximum, there was a gradual reduction in strength. The slopes of the load curves of connections decrease with the increasing cyclic load, but the slopes of the unloading curves remain almost constant, indicating loading stiﬀ ness degradation and a smaller deg-radation in the unloading stiﬀness. F igure 6 shows skeleton curves of the two specimens. Both specimens exhibited elastic, yielding, and hardening stages during the test. F or specimen T1, the steel tubes and concrete in this stage can be regarded as a whole, bearing the load together. Due to the cooperative working mechanism and deformation, T1 showed greater shear stiﬀ ness. In the yielding stage, the shear deformation increased with the increasing load. The concrete cracked and the web of the steel tube yielded. Afterwards, the slope of the curve declines. Then, the web of the steel tube entered the hardening phase until it buckled. After the web of the steel tube buckled, the shear resistance of the concrete compression strut gradually increased due to the conﬁ nement provided by the steel tube and diaphragms. Thus, the panel zone exhibited good ductility.F or specimen T2, the steel beam bore most of the load. In the elastic stage, the curve was almost vertical, indicating that the steel tubes and concrete were all in the elastic stage. Beyond the yielding point, the steel beam started losing stiﬀ -ness and the load increased more gently. After a large plastic deformation, the load reached its peak value and entered the stage of slow decline. Compared with that of T1, the load–displacement curve of T2 developed more moderately.S train Distribution The strain testing points were used to study the stress–strain distribution of the through-diaphragm connections. For ease of presentation, only the strains of panel zone are discussed below. The layout of the strain testing points in panel zone is presented in Fig. 7 . The points 24, 25, and 26 correspond to the corner area in three directions, and the points 27, 28, and 29correspond to the core. F igure 8 shows the shear-strain distributions in the panel zones of T1 and T2. It can be seen that the strain in the web of the steel tube was uniform prior to the yielding of the panel zone but was irregular after yielding. The points 25 and 29correspond to strain in the diagonal direction in the tube web.F ig. 5 H ysteretic loops of T1F ig. 6 L oad–displacement relationships of specimens. a T1 and b T2392 B. Rong et al.1 3Their values were higher than those in other directions. Themaximum value of T1 was about 15000, whereas that of T2 was only about 2500, indicating that T1 failed in the shear mode in the panel zone, but T2 did not.D uctility D uctility is deﬁ ned as the ability of the connection to undergo large amplitude deformations without any pronounced reduc-tion in strength. The ultimate displacement ductility μ is deﬁ ned in Eq. ( 1 ) [ 16 ], where Δu and Δy are the displacements of the beam end related to the ultimate and yielding strengths, respectively:(1)μ=Δu Δy. T he values of μ for T1 and T2 are 4.04 and 3.82, respec-tively. It can be seen that the connection in the shear fail-ure mode (T1) has greater ductility than that in bendingfailure mode (T2 in this paper and JD-1, -2, -3, -4 in Ref. [ 17 ]). Moreover, the ductility of through-diaphragm con-nections to pure steel beams is higher than that of con-nections to CFST columns and composite beams (all the specimens in Ref. [ 18 ]).S trength Degradation S trength degradation is deﬁ ned as the loss of strength under cyclic loading. Strength degradation can be evalu-ated by the strength deterioration λi , as calculated by Eq. (2)[19]: where n is the number of loading cycles at i times of Δy ; Pji is the peak load of the j t h cycle at i times of Δy ;and P y is the yielding load. As shown in Fig. 9 , the values of λi for T1 and T2 exhibit similar behavior prior to reaching the yield-ing strength. After that, T1 exhibits a more obvious strength improvement as well as strength degradation than T2. It can be seen that the shear failure mode can better develop the resistance of the through-diaphragm connection and that the bending failure mode will reduce the strength degradation of the connection. Overall, the strength degradation of connec-tions was not signiﬁcant in either failure mode.(2)λi =∑n j =1P ji ∕nP y, F ig. 7 L ayout of strain testing points in panel zoneF ig. 8 S train distribution in panel zone. a T1 and b T2393Experimental and Numerical Studies of Failure Modes and Load-Carrying Capacity of… 1 3N umerical StudyG eneral ABAQUS was used for the FE simulation of through-dia-phragm connections. The validated FE models were used for further parametric studies. The dimensions of connec-tions are introduced in Fig. 1 and Table 1 . Three-dimen-sional eight-node continuum elements (C3D8R) were used to simulate diﬀ erent components including inner concrete, steel tube, diaphragm, and steel beam in the connection. C3D8R is a general purpose 8-node linear brick element, with reduced integration and three translation degrees-of-freedom at each node. A numerical model with ﬁ ner mesh size was required to obtain an accurate FEA result.ner elements always lead to longer computation he steel tube, steel beam, and diaphragm adopt a bilin-E s , and Poisson’s ne the elastic behavior; yield ne the plastic behavior. oncrete damaged plasticity model (CDPM) [ 20 ] wasned by the steel tube, a stress–strain relation-21 ] was used.he input data of both the steel and concrete in the con-listed in Table 2 . A surface-to-surface interaction between concrete and steel tube with hard contact in normal direction and Cou-lomb friction in tangential direction was used in the FEM. The bottom plane of the column was constrained in three translation directions to simulate the pinned end condition. The top plane of the column and two translation degrees-of-freedom except for the axial direction of the column top was constrained. The section of the beam was coupling con-strained to a reference point, where the displacement con-trolling lateral loading was applied on. Figure 10 gives the FEM in ABAQUS .V alidation The FEM developed in this paper is validated by the experi-mental results of two specimens presented above. The vali-dation includes load–displacement curves, yielding and ulti-mate resistances of the specimens, and their failure modes.F ig. 9 C omparison of strength deterioration ratioF ig.10F EM of through-dia-phragm connection394 B. Rong et al.1 3Figure 11 shows a comparison between the stress distribu-tion and deformations predicted by FEM and our experimen-tal observations. It can be seen that the developed FEM can oﬀ er reasonable estimations on failure modes of both T1 and T2. In the FEM of T1, the shear deformation and the local buckling of the steel webs in shear failure mode are similar to those in the experiment. In the FEM of T2, the bending failure mode with a plastic hinge between the beam ﬂ ange and diaphragm resembles that in the experiment. F igure 12 shows a comparison of the load–displace-ment hysteretic curves and skeleton curves of the numerical analysis and experimental results. The stiﬀ ness and strength degradation in the predicted hysteretic loops are close to the experimental values of T1, and the predicted skeleton curve under monotonic loading is similar to the experimental curve of T2. Table 4 lists the yielding and ultimate loads of the two specimens in the experiment and FEA. The yield-ing lateral load ( P y ) and its corresponding displacement of each specimen are determined by the general yielding point method reported by Park et al. [ 22 ]. The above comparisons conﬁ rm that the developed FEMs are capable of predicting the behavior of specimens under both cyclic loading and monotonic loading. S ince cyclic loading sometimes causes diﬃ culty in com-putational convergence in FEA, monotonic loading is used in the following parametric studies. To verify the accuracy of this replacement, a simulation of T1 under monotonic load-ing is compared with that under cyclic loading. Figure 13 shows a comparison of two skeleton curves under two load-ing patterns, in which we can see that the change in load-ing patterns makes a little diﬀ erence in the load-carrying capacity or stiﬀ ness of the connection. The above veriﬁ ca-tion conﬁ rms that applying monotonic loading in subsequent parametric studies is practicable.P arametric StudyE ff ect of Tube Thickness in Panel ZoneT o investigate the eﬀ ect of tube thickness ( t ) in the panel zone on the failure mode and load-carrying capacity of the connection, the FEMs of T1 and T2 were used for an exten-sive parametric study. This parametric study included four levels for each model with varying thicknesses of the tube in the panel zone, i.e., t = 6, 8, 10, and 12 mm.F ig.11C omparison of thefailure mode results between experiment and FEM. a T1 in experiment, b T1 in ABAQUS, c T2 in experiment; d T2 in ABAQUS395Experimental and Numerical Studies of Failure Modes and Load-Carrying Capacity of… 1 3F igures 14 , 15 show comparisons of the load–displace-ment curves of T1 and T2, respectively. It can be seen that thas a little eﬀ ect on elastic stiﬀ ness. However, with increas-ing t , the yielding resistance of the connection is signiﬁ -cantly improved. It can be concluded that the steel tube in panel zone can provide an important contribution to the shear resistance of connection. The yielding and ultimate loads of all specimens are given in Fig. 16 . It should benoted that the increased range of the ultimate load-carryingcapacity gradually decreases with increasing t . F or T1, the thickness of tube increased by 33% (from 6 to 8 mm). As a result, the yielding load increased by 34% (from 114.6 to 154.2 kN), and the ultimate load increased by 19% (from 192.3 to 229.1 kN). The thickness of tube increased by 100% (from 6 to 12 mm), and the yielding load increased by 91% (from 114.6 to 218.4 kN). Contrary to expectations, the ultimate load increased by just 33% (from 192.3 to 256.4 kN). For T2, the thickness of tube increased by 33% (from 6 to 8 mm). As a result, the yielding load increased by 32%F ig.12C omparison of load–displacement curves. a T1 and b T2T able 4L oad-carrying capacity of specimensN ote: P t y is yielding load in experiment; P f y is yielding load in FEA;P t u is ultimate load in experiment; P f u is ultimate load in FEAS pecimen P t y (kN) P f y (kN) P t u (kN) P f u (kN) P f y ∕P ty P f u ∕P tuT1 120 114.6 198.8 192.3 0.95 0.97 T2 211.4 207.9 275.1 271.1 0.98 0.99 A verage––––0.97 0.98 F ig.13S keleton curves of T1 under monotonic and cyclic loadingsF ig. 14 P –Δ curves of T1 with diﬀ erent t values396B. Rong et al.1 3(from 129.5 to 171.1 kN), and the ultimate load increased by 21% (from 207.8 to 251.6 kN). The thickness of tube increased by 100% (from 6 to 12 mm) and the yielding load increased by 60% (from 129.5 to 207.9 kN). Similar to the previous situation, the ultimate load only increased by 30% (from 207.8 to 271.1 kN).T he possible reason for this result is that a large tube will cause the failure mode to change from shear failure in the panel zone to a plastic hinge on the beam ﬂ ange. The tube in the panel zone contributes little to the ultimate load when a plastic hinge appears. Figures 17 , 18 give the fail-ure modes of connections with diﬀ erent t values. It can be seen in Figs. 17 d and 18 d that the connections with stronger panel zone always fail in the beam ﬂ ange. In addition, a little shear deformation can be observed in the panel zone. However, obvious shear deformation occurs when the panel zone is weaker, as shown in Figs. 17 a , b and 18 a , b. When the resistances of the panel zone and beam are similar, thein Figs. 17c and 18c .ff ect of Diaphragm Thicknesso investigate the eﬀ ect of the thickness of diaphragm ( d ) d = 12, 14, and 16 mm. Figures 19 20 show a comparison of the load–displacement curves -ness, yield, and ultimate resistances of T1 and T2 are all extremely close, within deviations of only about 1%. The diaphragm only plays a force transfer role. As such, in the design codes for the load-carrying capacity of through-dia-phragm connections, the contribution of the diaphragm can be neglected. To facilitate construction, the diaphragms can be as thick as the beam ﬂ ange.C ritical Condition of Failure Modes T he occurrence of two failure modes, i.e., shear failure and bending failure, will be determined by the yield resistances of the panel zone and beam ﬂ ange. The shear yield resist-ance of the panel zone can be determined by the following equation, as developed by Nishiyama et al. [ 5 ]:where V py is the shear yield resistance of the panel zone; V psy is the shear yield resistance of the steel tube; and V pcy is the shear resistance of the concrete core.T he bending yield resistance of the steel beam can be determined by the following equation:(3)V py =V psy +V pcy, F ig. 15 P –Δ curves of T2 with diﬀ erent t valuesF ig.16P arametric study results. a T1 and b T2where M y is the bending yield resistance of the steel beam;f y is the yield strength of steel obtained from Table2 ; andW x is the elastic modulus of the beam section. (4)M y =f y ⋅W x ,In addition, the shear force in the panel zone of an exte-rior connection can be determined by the following:(5)V =M b1h b −V c =PL b H b −t bf −PLH,F ig.17F ailure modes of T1 with diﬀ erent t values. a t = 6 mm, b t = 8 mm, c t = 10 mm, and d t = 12 mmF ig.18F ailure modes of T2 with diﬀ erent t values. a t = 6 mm, b t = 8 mm, c t = 10 mm, and d t = 12 mmwhere Vc is reaction force; P is load on the beam end; L is distance from the beam end to the center of the column; L b is distance from the beam end to the column surface, H b is the depth of the beam; t bf is the thickness of the beam ﬂ ange; and H is the height of the column. Figure 21 shows the cal-culation diagram for the connection. T he lateral load P V corresponding to V py can be deter-mined as follows:T he lateral load P M corresponding to M y can be deter-mined as follows:F inally, eight calculation results for specimens in diﬀ erent failure modes are compared in Table 5 . It can be seen that the failure modes of connections are deter-mined by P V / P M of the connections. When P V / P M ≤ 0.9, the connection will fail in the shear mode. When 0.90 < P V / P M ≤ 1.23, the failure mode will be a combina-tion of the shear failure of the panel zone and the bending failure of the beam. When 1.23 < P V / P M , the connection will fail in the bending mode.(6)P V =V pyL bH b−t bf−L H.(7)P M =M yL b .F ig. 19 P –Δ curves of T1 with diﬀ erent d valuesF ig. 20 P –Δ curves of T2 with diﬀ erent d valuesPanel zone Steel beamCFSTcolumn M bM c2M c1V c =PL/HV c =PL/H N 0=P +NPLHNH bh bt bfM c1M bM c2V cV cNN 0PF ig.21P anel zone subjected to lateral force。

多孔介质有效导热系数的计算方法

第 30 卷第 3 期航空计算技术 2000 年 9 月
多孔介质有效导热系数的计算方法
潘宏亮
(西北工业大学航海工程学院 , 陕西西安 710072)
摘要 :高孔隙率多孔介质如泡沫陶瓷在新型多孔介质燃烧器技术中应用日益广泛 , 其重要传热特性参数 ——— 有效导热系数反映了两相流气、固相导热、对流和辐射的综合效应 , 对其研究尚非常缺乏。本文基于实验测定的温度分布 , 给出多孔介质有效导热系数的初始估值 , 用有限体积法求解二维控制方程 , 采用二维寻优搜索的办法 , 确定使测定点上测量与计算温度均方根误差为最小的径向与轴向有效导热系数 , 是一种逆计算方法。对球粒子颗粒床进行的有效性试验证明了方法的可行性。关键词 :有限体积法 ;导热系数 ;多孔介质 ;逆法中图分类号 :O242 , O551 文献标识码 :A
ond - O rder Accurate I mplicit Scheme .Computers and Fluids, (1974), Vo l.2 , pp.207 -209. [ 7] Tso tsas, E., and M artin , H .T hermal Conductivity of Packed Beds:A Review .Chem .Eng .Process , 1987 , V ol. 22, pp.19 -37 .
[ 2] Howell J.R ., Hall, M .J ., and Ellzey , J .L .Combustion of Hydro carbo n Fuels within Porous Iner t M edia . P rog .Energy Combustion Sci , 1996, Vol.22 , pp.121 145 .

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At the Magnel Laboratory for Concrete Research from Ghent University several research programmes have been conducted concerning massive concrete armour units for breakwaters, especially dealing with thermal stresses due to heat of hydration leading to cracks. A new method for the study of the inﬂuence of cracks in concrete structures on carbonation and chloride penetration has been developed [9–11]. This method consists of the use of notches by means of the positioning and removal afterwards of thin copper sheets inside the specimen in the fresh concrete. After few hours, the copper sheets were removed and artiﬁcial cracks were created in the concrete. The notched specimens have been subjected to a nonsteady state migration test. After the test, the specimens were axially split perpendicular to the notch and an AgNO3 solution was sprayed in order to measure the chloride penetration depth.
L. Marsavina et al. / Construction and Building Materials 23 (2009) 264–274
265
transport of ClÀ, K+, Na+ and OHÀ through saturated concrete during a migration test. Recently Martin-Perez et al. [7] proposed an algorithm based on modiﬁed Fick’s second law for numerical modelling of transport of chlorides, moisture, oxygen and heat convection through concrete. Wang et al. [8] proposed a mathematical model for the simulation of an electrochemical chloride removal process in order to predict the ionic mass transport associated with chloride ingress into concrete or hydrated cement paste from a saline environment. Up to now not many simulations were performed in order to determine the chloride penetration in cracked concrete geometries.
Available online at
Construction and Building Materials 23 (2009) 264–274
Construction and Building
MAt
Keywords: Chloride penetration; Crack; Concrete; Experimental determination; Numerical simulation
1. Introduction
Chloride-induced steel corrosion is worldwide one of the major deterioration problems for reinforced concrete structures. The high alkaline environment of good quality concrete forms a passive ﬁlm on the surface of the embedded steel which normally prevents the steel from further corroding. However, under chloride attack, the passive ﬁlm is disrupted or destroyed, and the steel spontaneously corrodes [1]. The volume of rust products is about four to six times larger than that of iron. This volume increase induces internal tensile stresses in the cover concrete, and when these stresses exceed the tensile strength of the concrete, the cover concrete is damaged by cracking, delamination and spalling. In addition to loss of cover concrete, a reinforced concrete member may suﬀer structural damage
a ‘‘Politehnica’’ University of Timisoara, Department Strength of Materials, Blvd. M. Viteazu, No. 1, Timisoara 300222, Romania b Ghent University, Magnel Laboratory for Concrete Research, Department of Structural Engineering, Technologiepark-Zwijnaarde 904, B-9052 Ghent, Belgium c RA Aquatim, Gh. Lazar Str., No. 11A, Timisoara 300081, Romania
Received 6 July 2007; received in revised form 16 December 2007; accepted 19 December 2007 Available online 21 February 2008
Abstract
One of the major causes of deterioration of reinforced concrete structures is chloride-induced corrosion of the reinforcing steel. This phenomenon is signiﬁcantly inﬂuenced by the presence of cracks. This paper presents the experimental and numerical results on the inﬂuence of cracks on chloride penetration in concrete structures. The experimental results were obtained with the non-steady state migration test, described in NT BUILD 492, using an electrical ﬁeld and artiﬁcial cracks. The numerical simulation results were obtained using a transient ﬁnite element analysis. A variable diﬀusion coeﬃcient with total chloride concentration was considered in the simulation. A good agreement between experimental and the numerical chloride penetration results was obtained. Ó 2008 Elsevier Ltd. All rights reserved.
* Corresponding author. Tel.: +40 256 403577; fax: +40 256 403523. E-mail address: msvina@mec.upt.ro (L. Marsavina).
0950-0618/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.conbuildmat.2007.12.015
Numerical studies were focused on developing a numerical algorithm in order to estimate the chloride penetration. Meijers [4] and Tang [5] used a mathematical model based on simulating free chloride penetration through the pore solution in concrete and calculating the distribution of the total chloride content in concrete. Truc et al. [6] present a numerical model, based on ﬁnite diﬀerence method and on the Nerst–Planck equation for the simulation of the