人人文库网 > 图纸下载 > 毕业设计 > YM3150E型滚齿机的控制系统的PLC改造设计【4张CAD图纸和说明书】

Modelling tool wear in cemented-carbide machining alloy.pdf

YM3150E型滚齿机的控制系统的PLC改造设计【4张CAD图纸和说明书】

资源目录

YM3150E型滚齿机的控制系统的PLC改造【4张CAD图纸和说明书】.rar

Modelling tool wear in cemented-carbide machining alloy.pdf---(点击预览)

PLC程序.dwg

状态转移图.dwg

电气原理图.dwg

电气接线图.dwg

压缩包内文档预览：

编号：14764931 类型：共享资源大小：1.29MB 格式：RAR 上传时间：2019-02-16 上传人：俊****计 IP属地：江苏

40
积分

版权申诉 word格式文档无特别注明外均可编辑修改；预览文档经过压缩，下载后原文更清晰！ 立即下载

关键词：: 4张CAD图纸和说明书 ym3150e 型滚齿机控制系统 plc 改造 cad 图纸以及说明书仿单

资源描述：

摘要 1

关键词 1

一.前言 2

1.1 PLC的国内外的状况 2

1.2 PLC的组成及特点 4

1.3 PLC的用途 5

1.4 PLC常用语言 6

1.5本课题的主要任务 6

二.整体方案的选择 7

2.1整体功能介绍 7

2.2控制要求 7

2.3电器说明…………………………………………………………………… .. 8

2.4保护装置……………………………………………………………………... 8

2.5电气原理图、电器接线图……………………………………………………. .8

三.硬件系统设计 9

3.1 PLC机型的选择步骤与原则 9

3.2 F1系列PLC的指令系统简介 10

3.3 YM3150E型滚齿机的硬件系统设计 11

四软件系统的设计 13

4.1 PLC程序设计的相关指令 13

4.2根据控制要求绘制梯形图 15

五结束语 16

小结…………………………………………………………………….......................................17

致谢…………………………………………………………………………………………….. 18

参考文献 ……………………………………………………………………………………. 19

附录1程序………………………………………………………………………………….... 20

附录2 外文文献翻译……………………………………………………………………… ..23

附录3外文文献原文……………………………………………………………….34

YM3150E型精密滚齿机的PLC改造

摘要

近年来，作为机电一体化重要技术的可编程序控制器(PLC)产品的集成度越来越高,工作速度越来越快,功能越来越强,使用越来越方便,特别是远程通信功能的实现,易于实现柔性加工和制造系统,使得PLC如虎添翼。本文简要的介绍了YM3150E型精密滚齿机的控制原理，并利用PLC对滚齿机进行改造，设计PLC控制系统，使滚齿机的控制更加方便。

关键词：滚齿机，控制系统，机电一体化，PLC

Reconstruction of YM3150E Precision Gear Hobbing Machine by PLC

Abstract

In recent years, as an important technology in Mechatronics, Programmable Logic Controller (PLC) products are more integrated, working faster and faster, more powerful in function, more and more convenient to use, especially in telecommunications function implementation, it is easy to implement flexible processing and manufacturing systems, makes the PLC even more powerful. This article briefly describes YM3150E Precision hobbing machine control principle, and to use PLC to reform of the hobbing machine, PLC control system designed to enable gear hobbing machine control more convenient.

Keyword：Hobbing Machine,Control System, Mechatronics, PLC

1.前言

现代科学技术的不断发展，极大地推动了不同学科的交叉与渗透，导致了工程领域的技术革命与改造。在机械工程领域，由于微电子技术和计算机技术的迅速发展及其向机械工业的渗透所形成的机电一体化，使机械工业的技术结构、产品机构、功能与构成、生产方式及管理体系发生了巨大变化，使工业生产由“机械电气化”迈入了“机电一体化”为特征的发展阶段。PLC作为机电一体化的一个重要的进程，在机械电气化的过程中起着很大的作用，现在还是这样，随着PLC本身的发展，它的应用范围越来越广，功能越来越强的。

可编程序控制器（programmable Logic Controller）是一种数字运算操作电子系统，专为在工业环境下应用而设计。它采用了可编程序的存储器，用来在其内部存储执行逻辑运算、顺序控制、定时、计数和算术运算等操作的指令，并通过数字的，模拟的输入和输出，控制各种类型的机械或生产过程。可编程序控制器及其有关的外围设备，都应按易于与工业控制系统形成一个整体、易于扩充其功能的原则设计。

1.1 PLC的国内外的状况

在工业生产过程中，大量的开关量顺序控制，它按照逻辑条件进行顺序动作，并按照逻辑关系进行连锁保护动作的控制，及大量离散量的数据采集。传统上，这些功能是通过气动或电气控制系统来实现的。1968年美国GM（通用汽车）公司提出取代继电气控制装置的要求，第二年，美国数字设备公司（DEC）研制出了基于集成电路和电子技术的控制装置，首次采用程序化的手段应用于电气控制，这就是第一代可编程序控制器，称Programmable ,是世界上公认的第一台PLC.

限于当时的元器件条件及计算机发展水平，早期的PLC主要由分立元件和中小规模集成电路组成，可以完成简单的逻辑控制及定时、计数功能。20世纪70年代初出现了微处理器。人们很快将其引入可编程控制器，使PLC增加了运算、数据传送及处理等功能，完成了真正具有计算机特征的工业控制装置。为了方便熟悉继电器、接触器系统的工程技术人员使用，可编程控制器采用和继电器电路图类似的梯形图作为主要编程语言，并将参加运算及处理的计算机存储元件都以继电器命名。此时的PLC为微机技术和继电器常规控制概念相结合的产物。个人计算机（简称PC）发展起来后，为了方便，也为了反映可编程控制器的功能特点，可编程序控制器定名为Programmable Logic Controller（PLC）。

20世纪70年代中末期，可编程控制器进入实用化发展阶段，计算机技术已全面引入可编程控制器中，使其功能发生了飞跃。更高的运算速度、超小型体积、更可靠的工业抗干扰设计、模拟量运算、PID功能及极高的性价比奠定了它在现代工业中的地位。20世纪80年代初，可编程控制器在先进工业国家中已获得广泛应用。这个时期可编程控制器发展的特点是大规模、高速度、高性能、产品系列化。这个阶段的另一个特点是世界上生产可编程控制器的国家日益增多，产量日益上升。这标志着可编程控制器已步入成熟阶段。

上世纪80年代至90年代中期，是PLC发展最快的时期，年增长率一直保持为30～40%。在这时期，PLC在处理模拟量能力、数字运算能力、人机接口能力和网络能力得到大幅度提高，PLC逐渐进入过程控制领域，在某些应用上取代了在过程控制领域处于统治地位的DCS系统。

20世纪末期，可编程控制器的发展特点是更加适应于现代工业的需要。从控制规模上来说，这个时期发展了大型机和超小型机；从控制能力上来说，诞生了各种各样的特殊功能单元，用于压力、温度、转速、位移等各式各样的控制场合；从产品的配套能力来说，生产了各种人机界面单元、通信单元，使应用可编程控制器的工业控制设备的配套更加容易。目前，可编程控制器在机械制造、石油化工、冶金钢铁、汽车、轻工业等领域的应用都得到了长足的发展。

我国可编程控制器的引进、应用、研制、生产是伴随着改革开放开始的。最初是在引进设备中大量使用了可编程控制器。接下来在各种企业的生产设备及产品中不断扩大了PLC的应用。目前，我国自己已可以生产中小型可编程控制器。上海东屋电气有限公司生产的CF系列、杭州机床电器厂生产的DKK及D系列、大连组合机床研究所生产的S系列、苏州电子计算机厂生产的YZ系列等多种产品已具备了一定的规模并在工业产品中获得了应用。此外，无锡华光公司、上海乡岛公司等中外合资企业也是我国比较著名的PLC生产厂家。可以预期，随着我国现代化进程的深入，PLC在我国将有更广阔的应用天地。

1.2 PLC的组成及特点

从结构上分，PLC分为固定式和组合式（模块式）两种。固定式PLC包括CPU板、I/O板、显示面板、内存块、电源等，这些元素组合成一个不可拆卸的整体。模块式PLC包括CPU模块、I/O模块、内存、电源模块、底板或机架，这些模块可以按照一定规则组合配置。

这里主要介绍一下它的CPU，CPU是PLC的核心，起神经中枢的作用，每套PLC至少有一个CPU，它按PLC的系统程序赋予的功能接收并存贮用户程序和数据，用扫描的方式采集由现场输入装置送来的状态或数据，并存入规定的寄存器中，同时，诊断电源和PLC内部电路的工作状态和编程过程中的语法错误等。进入运行后，从用户程序存贮器中逐条读取指令，经分析后再按指令规定的任务产生相应的控制信号，去指挥有关的控制电路。CPU主要由运算器、控制器、寄存器及实现它们之间联系的数据、控制及状态总线构成，CPU单元还包括外围芯片、总线接口及有关电路。内存主要用于存储程序及数据，是PLC不可缺少的组成单元。在使用者看来，不必要详细分析CPU的内部电路，但对各部分的工作机制还是应有足够的理解。CPU的控制器控制CPU工作，由它读取指令、解释指令及执行指令。但工作节奏由震荡信号控制。运算器用于进行数字或逻辑运算，在控制器指挥下工作。寄存器参与运算，并存储运算的中间结果，它也是在控制器指挥下工作。CPU速度和内存容量是PLC的重要参数，它们决定着PLC的工作速度，IO数量及软件容量等，因此限制着控制规模。

全部文件.png

PLC程序.gif

电气接线图.gif

电气原理图.gif

状态转移图.gif

字数统计.gif

内容简介：: Modelling tool wear in cemented-carbide machining alloy 718 J. Lorentzon?, N. Ja rvstra t Department of Technology, University West, Gustava Melins gata 2, Trollha ttan, Sweden a r t i c l e i n f o Article history: Received 3 January 2008 Received in revised form 28 February 2008 Accepted 2 March 2008 Available online 18 March 2008 Keywords: Tool wear FEM Inconel 718 Friction Modelling a b s t r a c t Tool wear is a problem in turning of nickel-based superalloys, and it is thus of great importance to understand and quantitatively predict tool wear and tool life. In this paper, an empirical tool wear model has been implemented in a commercial fi nite element (FE) code to predict tool wear. The tool geometry is incrementally updated in the FE chip formation simulation in order to capture the continuous evolution of wear profi le as pressure, temperature and relative velocities adapt to the change in geometry. Different friction and wear models have been analysed, as well as their impact on the predicted wear profi le assessed. Analyses have shown that a more advanced friction model than Coulomb friction is necessary in order to get accurate wear predictions, by drastically improving the accuracy in predicting velocity, thus having a dramatic impact on the simulated wear profi le. Excellent experimental agreement was achieved in wear simulation of cemented carbide tool machining alloy 718. fax: +46520223099. E-mail addresses: john.lorentzonhv.se, john.jl.lorentzon (J. Lorentzon). International Journal of Machine Tools W2. For Usuis model, a second set of parameters giving a different temperature dependency is also investigated; W3. Takayama and Muratas 17 wear rate model, including a functional dependence of absolute temperature; W4. Modifi ed Usui wear rate model, by adding an exponent on the relative velocity; W5. Vibration-adjusted Usui model, where a constant term is added to the relative velocity to account for vibrations, which are not included in the chip formation model. 1.1.2. Friction F1. Coulomb friction model, which states that the friction force is proportional to the contact pressure; F2. Shear friction model, which states that the friction force is a fraction of the equivalent stress; F3.Coulombfrictionmodelwithtwodifferentfriction coeffi cients, a reduced one around the tool tip and at a distance up on the rake face. 2. Tool wear model The tool wear model consists of a FE chip formation model and a wear model implemented as subroutines to calculate the wear rate at contact points, modifying the tool geometry accordingly. 2.1. Chip formation model The FE chip formation model was created using the commer- cial software MSC. Marc, which uses an updated Lagrange formulation. This means that the material is attached to the mesh, with periodic remeshing to avoid element distortion. The cutting process requires a coupled thermo-mechanical analysis, because mechanical work is converted into heat, causing thermal strains and infl uencing the material properties. Two types of thermal assumptions are commonly used for simulation of mechanical cutting, namely adiabatic heating and fully coupled thermalmechanical calculations. In this work a coupled, stag- gered, model has been used. This means that for each time increment the heat transfer analysis is made fi rst, followed by the stress analysis. The time increment was set to 1.5ms in all analyses. Quasi-static analyses were used, which means that the heat analysis is transient, while the mechanical analysis is static with inertial forces neglected. 2.1.1. Dimensions The dimensions of the workpiece used in the simulation model are 5mm length by 0.5mm height, and the tool used in the simulation model is 2mm long and 2mm high, see Fig. 1. The cutting-edgeradiuswassetto16mminagreementwith measurement (see Fig. 1), the clearance angle 61 and the rake angle 01; the feed was 0.1mm and the cutting speed was 0.75m/s. 2.1.2. Mesh The meshed workpiece can be seen in Fig. 1. The remeshing technique used was the advancing front quad. This mesh generator starts by creating elements along the boundary of the given outline boundary and mesh creation continues inward until the entire region has been meshed. The number of elements used was about 6000, with the minimum element size set at 2mm. As seen in Fig. 2a, fi ner mesh was used where the material separates around the tool tip. The tool was meshed with approximately 5000 elements, with minimum element size being 2mm. 2.1.3. Material properties Generally, the strain magnitude, the strain rate, and the temperature each have a strong infl uence on the material fl ow stress. Thus, it is necessary to capture these dependencies in the material model used, in order to correctly predict the chip formation. Here, neglecting a slight (about 10% between 1/s and ARTICLE IN PRESS Fig. 1. (a) Dimension of the chip formation model, scaled in mm. The start of the rake face is marked, as it will serve as a reference point in the wear profi les. (b) The measured cutting-edge shape, with the radius indicated; the measurement method is presented in Section 4.2. J. Lorentzon, N. Ja rvstra t / International Journal of Machine Tools a constant term is added to the relative velocity to account for vibrations not included in the chip formation model: dw dt A00snvrel 10e?B=T(8) 2.2.1. Wear model constants Constants for Usuis model (W1) were determined by Lor- entzon and Ja rvstra t 11 by calibrating machining simulations with measured wear rates: First, tool wear machining tests for the selected material were performed; then FE simulations were made under the same conditions; and fi nally the constants of the wear rate model were calculated by regression analysis, giving the constants B 8900 and A 1.82?10?12. This value of the B parameter is also used here, although the friction coeffi cient in the chip formation model differs because it is now calibrated with respect to feed force. For this reason, the A parameter was adjusted to give the same crater depth as in the experiments. The same methodology was used to calibrate A, D, A0and A00for the wear models W1, W2, W3, W4 and W5. The calibrated parameters are presented in Tables 2 and 3. The activation energy in W3 Eq. (6), E, was set to 75.35kJ/mol 25. 3. Analysis steps In a turning operation, a stationary condition with respect to temperature and forces will generally be reached almost im- mediately after the tool has penetrated the workpiece and subsequently the initial transient of the chip formation has been neglected in predictions of the tool wear progress. Instead, tool wear predictions are made on stationary chip formation condi- tions, and the fi rst step in tool wear predictions is therefore to calculate the stationary chip condition. Finally, the wear model is activated in the chip formation analysis and the progress of tool wear is calculated. 3.1. Chip formation In order to reach stationary conditions in FE chip formation simulations using the Lagrangian method, the entire object, on which chip formation simulation is to be performed, must be present and meshed from the beginning of the simulation. Consequentially a transient analysis for reaching steady-state conditions would be computationally prohibitive 26. Fortu- nately, by lowering the thermal capacity of the tool, it is possible to reach equilibrium faster; in our case this was obtained after about 1500 increments, see Fig. 5. The reason for this is that lowering the thermal capacity has the same effect as taking proportionally longer time increments in the thermal calculations compared to the mechanical increment, as can be seen in Eq. (5). Note that the left-hand side vanishes at steady state, while an increased Cpincreases the rate of change and correspondingly accelerates, reaching the steady-state condition: rCp qT qt k q2T qx2 q2T qy2 q2T qz2 ! (9) Here, T is the temperature, k is the thermal conductivity, r is the density, and Cpis the thermal capacity. 3.2. Tool wear The tool wear model consists of a FE chip formation model and a wear model implemented as subroutines calculating the wear rate at contact points, modifying the tool geometry accordingly. The wear rate is calculated using Usuis empirical wear model for every node of the tool in contact with the base material. In order ARTICLE IN PRESS Table 2 Wear model parameters for friction model F1 W1W2W3W4W5 A 1.25?10?12A 5.48?10?15D 1.02?10?9A0 5.42?10?13A00 1.08?10?15 B 8900B 2000E 75350B 8900B 8900 Table 3 Wear model parameters for W1 for the different friction models F1 (Coulomb)F2 (Shear)F3 (Adjusted) W1A 1.25?10?12A 1.52?10?12A 1.08?10?12 B 8900B 8900B 8900 J. Lorentzon, N. Ja rvstra t / International Journal of Machine Tools using more increments, however, would unnecessarily increase the computation time. The wear calculation corresponds to approximately 15s of dry machining, resulting in about 65mm fl ank wear land and about 5.3mm deep crater at rake face. Hence, the wear process is accelerated by about 10,000 times in the simulation model. 4. Experiments Turning experiments were conducted for calibration of friction and wear parameters and for comparison of the simulated and measured wear profi les. 4.1. Experimental conditions The turning experiments were carried out in a CNC lathe under dry cutting conditions. One cutting speed, vc, 45m/min and one feed rate, f, 0.1mm/rev were evaluated. The turning length machined during each experiment was 12mm. The experiment was conducted with 3 replicates. The workpiece was a bar of aged and forged Inconel 718 that was predrilled at its end face to achieve pipe geometry in order to accomplish near orthogonal cutting conditions in the turning operation. The workpiece had a 35.6mm outer diameter and a 29mm inner diameter. The tool used for these turning experiments was a triangular, uncoated cemented carbide turning insert with a cutting width of 16mm. Carbide classifi cation is in accordance with ISO standard N10-N30. 4.2. Measurements The cutting forces, the chip shape, the cutting edge radius, and the tool wear were all measured in these experiments. The cutting forces (the cutting force, Fc, the feed force, Ff, and the passive force, Fp) were measured for all samples using a three-component dynamometer(Kistler,Type9121),amulti-channelcharge amplifi er (Kistler, Type 5017B), and a data acquisition system. The chip shape was investigated for one sample using optical microscopy. For this purpose, the chips produced during each trial were collected, mounted onto specimen holders, grounded and polished. After this, the shape was investigated and the thickness was measured from the obtained images. For calibration and experimental verifi cation, wear profi les for the fl ank and the rake faces, as well as the cutting-edge radius, were measured for two cutting inserts. These measurements were performed at Toponova AB (www.toponova.se) using white-light interferometry, see e.g. 27 for a description. A schematic illustration of the tool in 3D with a cross-section where the worn and initial tool geometry was measured is presented in Fig. 7. 5. Results In this section, the impact of the wear and friction model on the simulated wear profi le is presented together with measured wear profi les.Simulatedtemperature,relativevelocityand contact pressure are also presented in order to emphasise and clarify the differences between the friction models. Finally, simulated cutting forces and chip thickness are compared with measurements. 5.1. Infl uence of wear model on wear profi le In this section, the Coulomb friction model (F1) is used and simulated wear profi les are compared with measured profi les using different wear equations. 5.1.1. Crater wear The simulated crater wear profi le using Usuis wear equation (W1) has a maximum crater depth in the upper region of the contact zone on the rake face, at about 200mm from the start of the rake face. This is in contrast to the measured wear profi le having a maximum crater depth close to the tool tip, at only about ARTICLE IN PRESS Fig. 5. Temperature history for two nodes in the tool, using thermal capacity acceleration. Fig. 6. Schematic illustration of the system for tool wear calculations. J. Lorentzon, N. Ja rvstra t / International Journal of Machine Tools (a) the rake face scaled to emphasise (length scale in mm), (b) the fl ank face scaled to emphasise (length scale in mm). J. Lorentzon, N. Ja rvstra t / International Journal of Machine Tools this phenomenon was presented by Desalvo and Shaw 28. The larger the friction coeffi cient, the larger this region of stationary, non-moving material and a large friction coeffi cient is necessary for good correlation between the simulated and measured feed force and contact length. However, by using a reduced friction coeffi cient around the tool tip, the velocity profi le changes dramatically in the stagnant region (from 0.01 to 0.14mm in Fig. 10). 5.3.2. Temperature The impact of friction model on predicted temperature in the tool in contact with the work piece can be seen in Fig. 11. The highest temperature is observed using the Coulomb friction model (F1), while the lowest is observed when using a Coulomb model with reduced friction in the tool tip region (F3). The difference in temperature prediction between the models is less than approximately 401C, with the shear model (F2) predicting temperatures in between the two others. The higher predicted temperature with constant Coulomb friction (F1) than the model with reduced friction (F3) may seem counterintuitive, as the heat generatedby contactfrictionislower.However,theheat generated by plastic deformation is instead correspondingly higher, as the relative motion is accommodated by material deformation, and in addition to this the material remains in this ARTICLE IN PRESS 350 300 250 200 150 100 50 0 -6 -6 -8 -10 -12 -14 -16 -80-60-40-200 -5-4-3-2-10 Measurement Measurement Initial profile Coulomb friction (F1) Shear friction (F2) Adjusted friction (F3) Fig. 9. Simulated wear profi les using different friction models (measured wear profi les and initial tool profi le as in Fig. 7); (a) the rake face scaled to emphasise (length scale in mm), (b) the fl ank face scaled to emphasise (length scale in mm). Fig. 10. Velocity profi le. J. Lorentzon, N. Ja rvstra t / International Journal of Machine Tools & Manufacture 48 (2008) 107210801078 zone for a longer time and consequently transports less of the generated heat away from this zone. For all models the temperature does not vary by more than about 1501C over the contact length and the shapes of the temperature profi les are very similar. It is, however, even more interesting that the temperature varies less than 251C over about three-quarter of the contact zone, from 0 to about 0.25mm in Fig. 11, the area where the tool wear is highest, as seen in Fig. 9. 5.3.3. Contact pressure Fig. 12 shows that the contact stress is highest at the tip of the tool, and that the maximum contact stress is very high, above 2.5GPa. Furthermore, the contact stress at the rake face stabilizes at two plateaus, one high, close to the tool tip, and one low, further up on the rake face. Interestingly enough, the contact pressure is very similar for all the different friction models, even though the contact lengths differ to some extent. 5.3.4. Cutting forces and chip thickness In this section, simulated cutting forces, chip thicknesses and contact lengths are presented and compared with measurements, see Table 4. The feed force (Ff) was used to calibrate the chip formation model. The cutting force (Fc), the chip thickness and the contact length were used for validation, showing less than 5% deviation. 6. Discussion It can be seen that quantitative agreement with measured wear is not possible using Coulomb friction (F1), regardless of which wear model is used. The largest discrepancy when using Usuis wear model (W1) is found in the region around the tool tip, between 0 and 100mm in Fig. 8. The simulated wear profi le can be changed somewhat by employing other wear models, but it seems impossible to achieve a correct position of the maximum crater depth, as shown by simulations employing wear models similar to the Usui model (W2, W4 and W5). This can be understood by studying the relative velocity graph (Fig. 10). This graph shows that simulations with traditional, constant coeffi cient, friction models (F1 and F2) predict zero velocity in the region where the highest measured crater depth is found. It would not help to vary pressure or temperature either, as Figs.11 and 12 show them to be very constant over the area of the wear peakthe temperature varies less than 251C. Thus, it is still too high up on the fl ank, and the shape of the crater is quite different from those measured. Applyingadiffusionwearmodel(W3)whichcompletely disregards the effect of velocity does allow a slightly better prediction of the position of maximum crater depth, but still too high, and the crater shape is very different from the measured ones. Having argued that it is insuffi cient to improve the wear model for quantitative and qualitative prediction of the wear profi le, instead, it is necessary to improve the simulation accuracy in the variables affecting wear. Of the three local variables in the wear model, the temperature and pressure distributions along the rake face seem impossible to infl uence enough to provide the observed wear shape through realistic model changes. The relative velocity, however, is dramatically infl uenced by friction. It seems reason- able to assume that the validity of Coulomb friction breaks down as the material approaches its yield limit. A very good agreement with the measured wear profi les was subsequently achieved by adding the reasonable physical assumption that the friction coeffi cient is lower i

温馨提示:
1: 本站所有资源如无特殊说明，都需要本地电脑安装OFFICE2007和PDF阅读器。图纸软件为CAD,CAXA,PROE,UG,SolidWorks等.压缩文件请下载最新的WinRAR软件解压。
2: 本站的文档不包含任何第三方提供的附件图纸等，如果需要附件，请联系上传者。文件的所有权益归上传用户所有。
3.本站RAR压缩包中若带图纸，网页内容里面会有图纸预览，若没有图纸预览就没有图纸。
4. 未经权益所有人同意不得将文件中的内容挪作商业或盈利用途。
5. 人人文库网仅提供信息存储空间，仅对用户上传内容的表现方式做保护处理，对用户上传分享的文档内容本身不做任何修改或编辑，并不能对任何下载内容负责。
6. 下载文件中如有侵权或不适当内容，请与我们联系，我们立即纠正。
7. 本站不保证下载资源的准确性、安全性和完整性, 同时也不承担用户因使用这些下载资源对自己和他人造成任何形式的伤害或损失。

人人文库网所有资源均是用户自行上传分享，仅供网友学习交流，未经上传用户书面授权，请勿作他用。