20米自動伸縮門設計【含CAD圖紙】,含CAD圖紙,20,自動,伸縮,設計,cad,圖紙
中期情況檢查表
學院名稱： 檢查日期：
學生姓名
專 業(yè)
指導教師
設計（論文）題目
20米自動伸縮門設計
工作進度情況
已經對自動伸縮門的總體結構有了一定的了解，對減速器的傳動方案做出選擇并進行了設計計算，采用了二級圓錐-圓柱齒輪減速器實現(xiàn)傳動?？刂葡到y(tǒng)已選擇為單片機控制系統(tǒng)，現(xiàn)正在對其原理圖進行設計并對對減速器設計進行驗證。
是否符合任務書要求進度
是
能否按期完成任務
能
工作態(tài)度情況
(態(tài)度、紀律、出勤、主動接受指導等)
對待畢業(yè)設計的態(tài)度較認真，按時完成每周任務，有問題及時向老師請教，與老師進行交流，并虛心接受老師的建議，遵守紀律，按時出勤。
質量
評價
(針對已完成的部分)
完成的畢業(yè)設計包括論文及圖紙，設計計算過程均為自己獨立完成。雖然存在一些小問題，但總體來說，設計質量還是可以的，確實認真做了工作。
存在問題和解決辦法
對圖紙的一些規(guī)定還不熟悉，存在一些問題，但通過一些制圖參考文件的參考能進行改進。對用單片機進行系統(tǒng)控制還無思路，可以去查閱相關資料來自我啟發(fā)。
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教學院長簽名
0
任 務 書
學 院（系）：
專 業(yè)：
學 生 姓 名：
學 號：
設計(設計)題目：
20米自動伸縮門設計
起迄日期:
指導教師:
教研室主任:
發(fā)任務書日期:
任 務 書
1．設計的背景：
本設計所設計的是20米自動伸縮門，隨著科技的提高與經濟的發(fā)展，各大、中、小型企事業(yè)單位（包括企業(yè)、研究單位、學校等）的大門口都用上了自動伸縮門。其樣式新穎、美觀、操作方便和功能更多（防盜、在上面顯示通告、遇故自動停止等）。目前，自動門設備大部分依靠進口，國內企業(yè)也開始試制自動門的機電梁系統(tǒng)。1994年中國建設部JG/T 3015.1、2—94《推拉自動門》、《平開自動門》發(fā)布，對規(guī)范自動門市場起到了一定的作用。隨著改革開放的加速和國際貿易的擴大，到目前包括日本、歐美等直接進口在內，已有近70個品牌的自動門機電梁設備進入中國市場。除推拉門和平開門外，弧形門和旋轉門等各種型號的自動門數量也日益增大，目前年需求量約有3萬臺，而且每年以30%的幅度遞增。
2．設計(設計)的內容和要求：
內容：
1）完成自動伸縮門減速器的設計，確定電動機的類型，計算傳動裝置的運動、動力參數，根據各參數，對各零件做強度校核。確定減速器類型。
2）完成鏈傳動設計，確定傳動方案，完成傳動的設計。
要求：
1) 有關文獻研讀
2) 設計說明書1份，20000字，打印裝訂成冊
3) 譯文5000字
4) 利用AutoCAD畫圖，圖紙數量折合成零號圖不少于3張。
3．主要參考文獻：
[1] 陳秀寧.主編.機械設計課程設計，浙江大學出版社，2007
[2] 杜白石.機械設計課程設計，西北農業(yè)大學，1993
[3] 龔桂義.機械設計課程設計指導書，高等教育出版社1992
[4] 邱宣懷.主編.機械設計(第四版)，高等教育出版社，1997
[5] 廖林清等.機械設計方法學，重慶大學出版社，1996.8
4．設計(設計)進度計劃(以周為單位)：
1周 設計調研
2周 設計調研
3周 總體方案設計
4周 運動動力參數計算
5周 零部件設計計算
6周 總裝配圖繪制
7周 總裝配圖繪制
8周 零部件圖繪制
9周 零部件圖繪制
10周 控制電路設計
11周 外文翻譯
12周 撰寫設計
13周 撰寫設計、準備答辯
教研室審查意見：
室主任簽名： 2014 年 2 月 21 日
學院審查意見：
教學院長簽名： 年 月 日
2
外文翻譯
Artificial hip joint
學生姓名
班 級
學院名稱
專業(yè)名稱
指導教師
1. Introduction
It has been recognised by a good number of researchers that the computation of the pressure distribution and contact area of artificial hip joints during daily activities can play a key role in predicting prosthetic implant wear [1], [2], [3] and [4]. The Hertzian contact theory has been considered to evaluate the contact parameters, namely the maximum contact pressure and contact area by using the finite element method [1] and [2]. Mak and his co-workers [1] studied the contact mechanics in ceramic-on-ceramic (CoC) hip implants subjected to micro-separation and it was shown that contact stress increased due to edge loading and it was mainly dependent on the magnitude of cup-liner separation, the radial clearance and the cup inclination angle [3] and [4]. In fact, Hertzian contact theory can captured slope and curvature trends associated with contact patch geometry subjected to the applied load to predict the contact dimensions accurately in edge-loaded ceramic-on-ceramic hips [5]. Although the finite element analysis is a popular approach for investigating contact mechanics, discrete element technique has also been employed to predict contact pressure in hip joints [6]. As computational instability can occur when the contact nodes move near the edges of the contact elements, a contact smoothing approach by applying Gregory patches was suggested [7]. Moreover, the contributions of individual muscles and the effect of different gait patterns on hip contact forces are of interest, which can be determined by using optimisation techniques and inverse dynamic analyses [8] and [9]. In addition, contact stress and local temperature at the contact region of dry-sliding couples during wear tests of CoC femoral heads can experimentally be assessed by applying fluorescence microprobe spectroscopy [10]. The contact pressure distribution on the joint bearing surfaces can be used to determine the heat generated by friction and the volumetric wear of artificial hip joints [11] and [12]. Artificial hip joint moment due to friction and the kinetics of hip implant components may cause prosthetic implant components to loosen, which is one of the main causes of failure of hip replacements. Knee and hip joints' moment values during stair up and sit-to-stand motions can be evaluated computationally [13]. The effect of both body-weight-support level and walking speed was investigated on mean peak internal joint moments at ankle, knee and hip [14]. However, in-vivo study of the friction moments acting on the hip demands more research in order to assess whether those findings could be generalised was carried out [15].
The hypothesis of the present study is that friction-induced vibration and stick/slip friction could affect maximum contact pressure and moment of artificial hip joints. This desideratum is achieved by developing a multibody dynamic model that is able to cope with the usual difficulties of available models due to the presence of muscles, tendons and ligaments, proposing a simple dynamic body diagram of hip implant. For this purpose, a cross section through the interface of ball, stem and lateral soft and stiff tissues is considered to provide the free body diagram of the hip joint. In this approach, the ball is moving, while the cup is considered to be stationary. Furthermore, the multibody dynamic motion of the ball is formulated, taking the friction-induced vibration and the contact forces developed during the interaction with cup surface. In this study, the model utilises available information of forces acting at the ball centre, as well as angular rotation of the ball as functions of time during a normal walking cycle. Since the rotation angle of the femoral head and their first and second derivatives are known, the equation of angular momentum could be solved to compute external joint moment acting at the ball centre. The nonlinear governing equations of motion are solved by employing the adaptive Runge–Kutta–Fehlberg method, which allows for the discretisation of the time interval of interest. The influence of initial position of ball with respect to cup centre on both maximum contact pressure and the corresponding ball trajectory of hip implants during a normal walking cycle are investigated. Moreover, the effects of clearance size, initial conditions and friction on the system dynamic response are analysed and discussed throughout this work.
2. Multibody dynamic model of the artificial hip joint
The multibody dynamic model originaly proposed by Askari et al. [16] has been considered here to address the problem of evaluating the contact pressure and moment of hip implants. A cross section A-A of a generic configuration of a hip joint is depicted in the diagram of in Fig. 1, which represents a total hip replacement. Fig. 1 also shows the head and cup placed inside of the pelvis and separated from stem and neck. The forces developed along the interface of the ball and stem are considered to act in such a way that leads to a reaction moment, M. This moment can be determined by satisfying the angular motion of the ball centre during a walking cycle. The available data reported by Bergmann et al. [17] is used to define the forces that act at the ball centre. This data was experimentally obtained by employing a force transducer located inside the hip neck of a live patient. The information provided deals with the angular rotation and forces developed at the hip joint. Thus, the necessary angular velocities and accelerations can be obtained by time differentiating the angular rotation. Besides the 3D nature of the global motion of the hip joint, in the present work a simple 2D approach is presented, which takes into account the most significant hip action, i.e. the flexion-extension motion. With regard to Fig. 2 the translational and rotational equation of motion of the head, for both free flight mode and contact mode, can be written by employing the Newton–Euler's equations [18] and [19], yielding
equation(1)
∑MOk=Iθ¨k,∑MO={Mk?(Rj)n×FPjtδ>0Mkδ≤0
equation(2)
∑FX=mx¨,∑FX={fx+(FPjt+FPjn)?iδ>0fxδ≤0
equation(3)
∑FY=my¨,∑FY={fy+(FPjn+FPjt)?j?mgδ>0fyδ≤0
where
FPjn and
FPjt denote the normal and tangential contact forces developed during the contact between the ball and cup, as it is represented in the diagram of Fig. 3. In Eqs. (1), (2) and (3), x, y and θ are the generalised coordinates used to define the system's configuration. In turn, variable m and I are the mass and moment of inertia of ball, respectively. The external generalised forces are denoted by fx, fy and M and they act at the centre of the ball as it is shown in Fig. 3. The gravitational acceleration is represented by parameter g, Rj is the ball radius and δ represents relative penetration depth between the ball and cup surfaces.
Fig. 1.
A schematic of the artificial hip implant with the cross section A-A (Left figure), and the head and cup separated from the neck and stem through the cross section A-A (Right figure).
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Fig. 2.
A schematic of the head and cup interaction observed in the Sagittal plane.
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Fig. 3.
Free body diagram of ball and corresponding external, internal and body forces and moment.
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The
Download as PowerPoint slideThe penetration depth can be expressed as [20]
equation(4)
δ=r?(Rb?Rj)δ=r?(Rb?Rj)
in which Rb denotes the cup radius and (Rb?Rj) represents the joint radial clearance, which is a parameter specified by user.
In the present study, the cup is assumed to be stationary, while the head describes the global motion. With regard to Fig. 2, it can be observed that Oj and Ob denote the head and cup centres, respectively. While Pj and Pb represent the contact points on the head and cup, respectively. The magnitude and orientation of the clearance vector are denoted by r and α, respectively. In general, r and α can be expressed as functions of the generalised coordinates used to describe the configuration of multibody mechanical system. The normal and tangential unit vectors at the contact point can be written as
equation(5)
n=cosαi+sinαj
equation(6)
t=?sinαi+cosαj
In order to compute the normal contact and tangential forces, it is first necessary to evaluate the relative tangential and normal velocities at the contact points, which can be obtained as follows: [17]
equation(7)
vpj/pb=r?n+(rα?+Rjωj)t=vnn+vtt
where vnvn and vtvt are module of the normal and tangential velocities, respectively. Thus, Eqs. (2) and (3) can be re-written as follows:
equation(8)
[m00m][x¨y¨]=[∑FX∑FY]
Using now the concept of the state space representation, the second order equations of motion (8), can be expressed as a first order equation set as
equation(9)
z?=H(z)
where
z=[z1z2z3z4]=[xyx?y?] and H(z)H(z) is expressed as follows:
equation(10)
z?=[z?1z?2z?3z?4]=[z3z4∑FX(z)∑FY(z)]
It must be mentioned that the r, α and their time derivatives can be obtained with respect to state space parameters as follows:
equation(11)
α=atan(z2z1)
equation(12)
r=z12+z22
equation(13)
α?=?z2z3+z1z4z12+z22z12
equation(14)
r?=z1z3+z2z4z12+z22
It is known that the evaluation of the contact forces developed during an impact event plays a crucial role in the dynamic analysis of mechanical systems [21], [22] and [23]. The contact forces must be computed by using a suitable constitutive law that takes into account material properties of the contacting bodies, the geometric characteristics of impacting surfaces and impact velocity. Additionally, the numerical approach for the calculation of the contact forces should be stable in order to allow for the integration of the mechanical systems equations of motion [24]. Different constitutive laws are suggested in the literature, being one of the more prominent proposed by Hertz [25]. However, this law is purely elastic in nature and cannot explain the energy loss during the impact process. Thus, Lankarani and Nikravesh [26] overcame this difficulty by separating the contact force into elastic and dissipative components as
equation(15)
Fpjn=(Kδn+Dδ?)n
Regarding Lankarani and Nikravesh model, normal contact force on the head is expressed as
equation(16)
Fpjn=?Kδ3/2(1+3(1?ce2)4δ?δ?(?))n
where
δ? and
δ?(?) are the relative penetration velocity and the initial impact velocity, respectively, and ce is the coefficient of restitution. The generalised stiffness parameter K depends on the geometry and physical properties of the contacting surfaces, which for two internal spherical contacting bodies with radii Ri and Rj can be expressed as [25]
equation(17)
K=43(σi+σj)(RiRjRi?Rj)2
in which the material properties σiσi and σjσj are given by
equation(18)
σz=1?υz2Ez
At this stage, it must be said the use of Eq. (15) is limited by Love's criterion, that is, it is only valid for impact velocities lower than the propagation velocity of elastic waves across the solids [27].
It is known that the way in which the friction phenomena are modelled, plays a key role in the systems behaviour [28]. In the present study, the tangential friction force are evaluated by using a modified Coulomb friction law, which can be expressed as [29] and [30]
equation(19)
Fpjt=?μ(vt)||Fpjn||vt|vt|t
The friction force is described in the sense of Coulomb's approach, and is proportional to the magnitude of the normal force developed at the contact points, where the ratio is the coefficient of friction, μ, which is dependent on the relative tangential velocity. The model considered in reference [17] is employed here for the purpose of evaluating the coefficient of friction, which can be written as
equation(20)
μ(vt)={(?cfv02(|vt|?v0)2+cf)sgn(vt),|vt|
Fig. 4.
Stribeck characteristic for dry friction.
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Normal and tangential forces described above are present if the system is
in contact situation, which means detecting impact or contact is one important step. Moreover, to compute the contact force, the initial impact velocity has to be calculated as an initial condition for following regimes, which could be in contact or in free flight, the following condition should be checked during the solution process by progressing time. Therefore, a contact event is detected when the following condition is verified
equation(21)
δ(ti)<0,δ(ti+1)>0
Indeed, the precise instant i