3S_Formability of Aluminum Sheet - 2013

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sus201不锈钢材质报告

sus201不锈钢材质报告English Answer:SUS201 Stainless Steel Material Report.SUS201 is a grade of austenitic stainless steel that is widely used in various applications due to its combination of strength, corrosion resistance, and affordability. This material report provides a comprehensive overview of the properties, composition, and applications of SUS201 stainless steel.Composition:SUS201 stainless steel is primarily composed of iron, chromium, nickel, and manganese. The typical chemical composition of SUS201 is as follows:Carbon (C): 0.15%。

Chromium (Cr): 16-18%。

Nickel (Ni): 3.5-5.5%。

Manganese (Mn): 6-8%。

Silicon (Si): 1.0%。

Copper (Cu): 0.25%。

Phosphorus (P): 0.045%。

Sulfur (S): 0.03%。

Properties:SUS201 stainless steel possesses a range of desirable properties, including:High strength and hardness.Excellent corrosion resistance.Good weldability and formability.Non-magnetic.Heat resistance up to 870°C (1600°F)。

A356.2铝合金轮毂拉伸性能及断口分析

AT &M
（上接第41页）铆钉用量在5 000~8 000颗不等，平均按2元/颗计算，仅铆钉材料成本一项就1万多元，这也是制约铝合金客车普及的一个重要因素。
车身蒙皮的连接以粘接为主[4]，在需要局部加强的部位配以自冲铆，即粘接-自冲铆复合工艺[5]，既满足表面要求又保证了连接强度。
5 总结
观结构细化，二次枝晶臂间距较小；后凝固部分随着保温时间延长，晶粒呈长大趋势，二次枝晶臂间距尺寸增加。 3.2 拉伸试验
轮毂上、下轮缘在同一圆周方向上各取2个样品，轮辐6个样品分别取自不同的辐条。相同部位不同样品的各力学性能参数分别取平均值得出轮毂上不同取样位置的抗拉强度、屈服强度和伸长率的对比关系，见图3。
sintered Al-Si-Mg matrix composites［J］. Journal of Materials Engineering and Performance，1997，6（6）: 785. ［3］Atxaga G, Pelayo A, Iri sar ri A M. Effect of microstructure on fatigue behavior of cast Al7-Si-Mg alloy［J］. Mater Sci Tech ，2001, 17: 446. ［4］Groteke D E, Designs Q C, et al. Influence of SNIF treatment on characteristics of aluminum foundry alloys ［J］. AFS Transactions，1985， 181: 953- 960. ［5］黄良余，张少宗. 铝合金精炼理论要点和工艺原则［J］ . 特种铸造及有色合金，1998，（2）: 40- 42. ［6］冉广，周敬恩，王永芳. 铸造A356铝合金的拉伸性能及其断口分析［J］. 稀有金属材料与工程，2006，35 （10）:1620-1624. ［7］张新梅，郝丽华，蒋大鸣，庞振民. A1-Mg-Si 合金拉伸断口研究［J］. 材料工程, 1996, 5: 35-36，15.

铝合金板阵列微结构零件电磁冲击液压成形研究

第16卷第3期精密成形工程2024年3月JOURNAL OF NETSHAPE FORMING ENGINEERING123铝合金板阵列微结构零件电磁冲击液压成形研究颜子钦1a，赵鹏1a，朱玉德2，阳光1b，王瀚鹏1b，徐勇3，崔晓辉1b,2*（1.中南大学 a.机电工程学院 b.轻合金研究院，长沙 410083；2.恩普赛技术有限公司，湖北襄阳 441021；3.中国科学院金属研究所，沈阳 110016）摘要：目的解决室温条件下因铝合金塑性流动不均而导致的零件开裂和尺寸偏差等问题。

方法利用高速冲击提高材料成形极限以及流体均匀载荷精确控形的优势，提出了电磁冲击液压工艺并实现了铝合金阵列结构零件的成形，采用实验手段研究了放电电压和放电次数对零件贴模精度和厚度分布的影响。

结果随着放电电压的增大，零件的成形深度增大。

在单次放电8 kV下，板料最大成形深度达到模具深度的97%，连续3次放电8 kV后，零件通道填充率达到89.7%。

建立了与物理实验模型一致的电磁-流体-结构的多物理场耦合仿真模型，发现冲击液体对板料施加的瞬态压强超过200 MPa，板料最大变形速度达到40.5 m/s。

模拟得到的板料变形轮廓与实验结果一致，证明了多物理场耦合仿真模型的准确性。

结论电磁冲击液压成形是一种新型的高速成形方法，能够实现铝合金阵列微结构零件的精确制造，为提高复杂薄壁难变形构件的成形性能和精度提供了新的技术手段。

关键词：电磁冲击液压；阵列微结构；高速率成形；多物理场耦合仿真；流固耦合DOI：10.3969/j.issn.1674-6457.2024.03.012中图分类号：TG391 文献标志码：A 文章编号：1674-6457(2024)03-0123-08Experimental Study on Electromagnetic Impact Hydraulic Forming ofAluminum Alloy Sheet Array Micro-structure PartsYAN Ziqin1a, ZHAO Peng1a, ZHU Yude2, YANG Guang1b, WANG Hanpeng1b, XU Yong3, CUI Xiaohui1b,2*(1.a. College of Mechanical and Electrical Engineering, b. Light Alloy research Institute, Central South University, Changsha410083, China; 2. EMPuls Technology Co., Ltd., Hubei Xiangyang 441021, China; 3. Institute of Metal Research,Chinese Academy of Sciences, Shenyang 110016, China)ABSTRACT: The work aims to solve the problems such as cracks and dimensional deviations caused by uneven plastic flow of aluminum alloy material at room temperature. Using the advantages of high speed impact to improve the forming limit of mate-rials and precise shape control of uniform fluid load, an electromagnetic impact hydraulic technology was put forward and the forming of aluminum alloy array structure parts was realized. The effects of discharge voltage and discharge times on the accu-racy and thickness distribution of the parts were studied through experiments. With the increase of discharge voltage, the form-ing depth of the part increased. Under discharge voltage of 8 kV, the maximum forming depth of sheet reached 97% of the die收稿日期：2024-01-08Received：2024-01-08基金项目：国家自然科学基金（52275394）；中南大学高性能复杂制造国家重点实验室项目（ZZYJKT2020-02）Fund：The National Natural Science Foundation of China (52275394)；The Project of State Key Laboratory of High Performance Complex Manufacturing, Central South University (ZZYJKT2020-02)引文格式：颜子钦，赵鹏，朱玉德, 等. 铝合金板阵列微结构零件电磁冲击液压成形研究[J]. 精密成形工程, 2024, 16(3): 123-130. YAN Ziqin, ZHAO Peng, ZHU Yude, et al. Experimental Study on Electromagnetic Impact Hydraulic Forming of Aluminum Alloy Sheet Array Micro-structure Parts[J]. Journal of Netshape Forming Engineering, 2024, 16(3): 123-130.*通信作者（Corresponding author）124精密成形工程 2024年3月depth, and the filling rate of parts channel reached 89.7% after three consecutive discharges of 8 kV. A multi-physical coupling simulation model of electromagnetic-fluid-structure was established, which was consistent with the physical experiment model.It was found that the transient pressure exerted by the impact liquid on the plate exceeded 200 MPa, and the maximum sheet de-formation velocity reached 40.5 m/s. The deformation profile of sheet obtained by simulation was consistent with the experi-mental results, which proved the accuracy of the multi-physics coupling simulation model. Electromagnetic impact hydraulic forming is a new high-speed forming method, which can realize the precise manufacturing of aluminum alloy array mi-cro-structure parts, and provide a new technical means to improve the formability and accuracy of complex thin-wall refractory components.KEY WORDS: electromagnetic impact hydraulic forming; micro array-structure; high speed impact forming; multi-physics coupling simulation; fluid-structure interaction铝合金具有较高的比强度和较好的耐腐蚀性，它制成的阵列微结构零件被广泛应用于交通运输行业[1]。

损伤初始及扩展

Damage initiation criterion and damage evolution responseAbaqus/Standard and Abaqus/Explicit offer a general capability for predicting the onset of failure and a capability for modeling progressive damage and failure of ductile metals. In the most general case this requires the specification of the following:•the undamaged elastic-plastic response of the material (“Classical metal plasticity,” Section 20.2.1);• a damage initiation criterion (“Damage initiation for ductile metals,” Section21.2.2); and• a damage evolution response, including a choice of element removal (“Damage evolution and element removal for ductile metals,” Section 21.2.3). Damage initiation criterion•Damage initiation criteria for the fracture of metals, including ductile and shear criteria.•Damage initiation criteria for the necking instability of sheet metal. These include forming limit diagrams (FLD, FLSD, and MSFLD) intended to assess the formability of sheet metal and the Marciniak-Kuczynski (M-K) criterion(available only in Abaqus/Explicit) to numerically predict necking instabilityin sheet metal taking into account the deformation history.More than one damage initiation criterion can be specified for a given material. If multiple damage initiation criteria are specified for the same material, they are treated independently.Damage evolutionThe damage evolution law describes the rate of degradation of the material stiffness once the corresponding initiation criterion has been reached. For damage in ductile metals Abaqus assumes that the degradation of the stiffness associated with each active failure mechanism can be modeled using a scalar damage variable, (), where represents the set of active mechanisms. At any given time during the analysis the stress tensor in the material is given by the scalar damage equationwhere D is the overall damage variable and is the effective (or undamaged) stress tensor computed in the current increment. are the stresses that would existin the material in the absence of damage. The material has lost its load-carrying capacity when . By default, an element is removed from the mesh if all of the section points at any one integration location have lost their load-carrying capacity.Input File Usage: U se the following option immediately after thecorresponding *DAMAGE INITIATION option to specifythe damage evolution behavior:*DAMAGE EVOLUTIONAbaqus/CAE Usage: P roperty module: material editor: Mechanical Damage forDuctile Metals criterion: Suboptions DamageEvolution1. Ductile criterionThe ductile criterion is a phenomenological model for predicting the onset of damage due to nucleation, growth, and coalescence of voids. The model assumes that the equivalent plastic strain at the onset of damage, , is a function of stress triaxiality and strain rate:where is the stress triaxiality, p is the pressure stress, q is the Mises equivalent stress, and is the equivalent plastic strain rate. The criterion for damage initiation is met when the following condition is satisfied:where is a state variable that increases monotonically with plastic deformation. At each increment during the analysis the incremental increase in is computed as2. Johnson-Cook criterionThe Johnson-Cook criterion (available only in Abaqus/Explicit) is a special case of the ductile criterion in which the equivalent plastic strain at the onset of damage, , is assumed to be of the formwhere – are failure parameters and is the reference strain rate. This expression differs from the original formula published by Johnson and Cook (1985) in the sign of the parameter . This difference is motivated by the fact that most materials experience a decrease in with increasing stress triaxiality;therefore, in the above expression will usually take positive values. is the nondimensional temperature defined as3. Shear criterionThe shear criterion is a phenomenological model for predicting the onset of damage due to shear band localization. The model assumes that the equivalent plastic strain at the onset of damage, , is a function of the shear stress ratio and strain rate:Here is the shear stress ratio, is the maximum shear stress, and is a material parameter. A typical value of for aluminum is = 0.3 (Hooputra et al., 2004). The criterion for damage initiation is met when the following condition is satisfied:4. Forming limit diagram (FLD) criterionThe forming limit diagram (FLD) is a useful concept introduced by Keeler and Backofen (1964) to determine the amount of deformation that a material can withstand prior to the onset of necking instability. The maximum strains that a sheet material can sustain prior to the onset of necking are referred to as the forming limit strains. A FLD is a plot of the forming limit strains in the space of principal (in-plane) logarithmic strains. In the discussion thatfollows major and minor limit strains refer to the maximum and minimum values of the in-plane principal limit strains, respectively. The major limit strain is usually represented on the vertical axis and the minor strain on the horizontal axis, as illustrated in Figure 21.2.2–1. The line connecting the states at which deformation becomes unstable is referred to as the forming limit curve (FLC). The FLC gives a sense of the formability of a sheet of material. Strains computednumerically by Abaqus can be compared to a FLC to determine the feasibility of the forming process under analysis.Figure 21.2.2–1 Forming limit diagram (FLD).The FLD damage initiation criterion requires the specification of the FLC in tabular form by giving the major principal strain at damage initiation as a tabular function of the minor principal strain and, optionally, temperature and predefinedfield variables, . The damage initiation criterion for the FLD is given by the condition , where the variable is a function of the current deformation state and is defined as the ratio of the current major principal strain, , to the major limit strain on the FLC evaluated at the current values of the minor principal strain, ; temperature, ; and predefined field variables, :For example, for the deformation state given by point A in Figure 21.2.2–1 thedamage initiation criterion is evaluated as5. Forming limit stress diagram (FLSD) criterionWhen strain-based FLCs are converted into stress-based FLCs, the resulting stress-based curves have been shown to be minimally affected by changes to the strain path (Stoughton, 2000); that is, different strain-based FLCs, corresponding to different strain paths, are mapped onto a single stress-based FLC. This property makes forming limit stress diagrams (FLSDs) an attractive alternative to FLDs for the prediction of necking instability under arbitrary loading. However, the apparent independence of the stress-based limit curves on the strain path maysimply reflect the small sensitivity of the yield stress to changes in plastic deformation. This topic is still under discussion in the research community.A FLSD is the stress counterpart of the FLD, with the major and minor principal in-plane stresses corresponding to the onset of necking localization plotted on the vertical and horizontal axes, respectively.Damage evolutionFigure 21.2.3–1 illustrates the characteristic stress-strain behavior of a material undergoing damage. In the context of an elastic-plastic material with isotropic hardening, the damage manifests itself in two forms: softening of the yield stress and degradation of the elasticity. The solid curve in the figure represents the damaged stress-strain response, while the dashed curve is the response in the absence of damage. As discussed later, the damaged response depends on the element dimensions such that mesh dependency of the results is minimized.Figure 21.2.3–1 Stress-strain curve with progressive damage degradation.In the figure and are the yield stress and equivalent plastic strain at the onsetof damage, and is the equivalent plastic strain at failure; that is, when the overall damage variable reaches the value . The overall damage variable, D, captures the combined effect of all active damage mechanisms and is computed in terms of the individual damage variables, , as discussed later in this section (see “Evaluating overall damage when multiple criteria are active”).The value of the equivalent plastic strain at failure, , depends on the characteristic length of the element and cannot be used as a material parameter for the specification of the damage evolution law. Instead, the damage evolution law is specified in terms of equivalent plastic displacement, , or in terms of fracture energy dissipation, ; these concepts are defined next.。

欧洲标准EN-10143连续热浸镀层薄钢板、钢带尺寸、外形允许偏差中文版

连续热浸镀层薄钢板、钢带尺寸、外形允许偏差EN10143:19931范围1.1该欧洲标准规定了连续热浸金属涂层扁钢材的尺寸和外形的偏差（所有宽度或从扁钢产品上分切下来的薄板）厚度W3.0mm的用于冷成型和结构用途的低碳钢钢材。

该厚度是含镀层的最终的交货产品厚度。

1.2该欧洲标准适用于所有热浸金属涂层的扁钢产品，如：一锌或铁锌合金（见EN10142和EN10147）；—锌铝合金（见EN10215正在起草中）；—铝锌合金（见EN10214,正在起草中）；—硅铝合金（见EURONORM 154,正在起草中）；一铅合金（见EURONORM 153）。

只要在每次业务中没有不同的或附加的技术交货条件在定货时就不用再另行协商了。

1.3该欧洲标准不适用于：—冷轧的或热轧的无涂层扁钢材（见EN10131和EN10051 ）；—电镀涂层的宽扁钢产品（见EN10152,举例说明）标准规定的钢材公差。

2参考标准本欧洲标准按日期和不按日期引用了其它的出版物的条款和参考资料。

这些引用的参考资料列在文本的恰当的地方和后面列出的刊物中。

对于过时的参考资料后来又经修改的，或修改了的任何适用于欧洲标准的出版物只能使用他们修改后的或修订后的内容。

对于未注明日期的参考资料请见援引了这些参考物的最新版本。

EN10020钢材的定义和钢级别分类EN10079钢材的定义EN10142连续热镀锌低碳冷轧钢带和薄板供货技术条件EN10147连续热浸镀锌结构钢板和钢带的交货条件EN10214连续热浸镀锌铝（ZA）钢板和钢带的交货条件EN10215连续热浸镀铝锌（AZ）钢板和钢带的交货条件EURONORM 153：19802热浸铅锡（铅合金）涂层冷还原碳轧钢扁钢材的商品级和冲压级—交货条件EURONORM 153：19802热浸铝硅涂层轧钢扁钢材的一般冲压级一交货条件3定义除了在EN10020和EN10079中规定的定义外，下面的定义也可以适用于本欧洲标准。

Al-Si镀层热成形零件表面颜色差异性研究

Al-Si镀层热成形零件表面颜色差异性研究单明东;夏益新;王娜;刘海涛;曹奇;张丹荣【摘要】目的研究加热温度、加热时间等工艺参数对Al-Si镀层材料在热成形过程中存在的表面颜色差异、镀层厚度和扩散层厚度的影响规律,及影响零件表面颜色差异的主要原因.方法在不同加热时间及加热温度条件下,对厚度为1.0 mm的新日铁Al-Si镀层材料进行热冲压试验,测量热成形零件的镀层厚度和扩散层厚度,并对典型不同颜色零件表面进行SEM及EDS分析研究.结果 Al-Si镀层热成形零件表面颜色与加热温度和加热时间存在较好的对应关系,同时镀层厚度及扩散层厚度随着加热时间的增加及加热温度的提高而增大,Al-Si镀层热成形零件表面的颜色与镀层中不同铁氧化物的混合比例存在较好的对应一致性.结论 Al-Si镀层热成形零件表面颜色的状态可以间接反应镀层厚度及扩散层厚度.【期刊名称】《精密成形工程》【年(卷),期】2017(009)006【总页数】5页(P99-103)【关键词】Al-Si镀层;锌基镀层;热成形;表面颜色差异【作者】单明东;夏益新;王娜;刘海涛;曹奇;张丹荣【作者单位】上海宝钢高新技术零部件有限公司,上海 201908;上海宝钢高新技术零部件有限公司,上海 201908;上海宝钢高新技术零部件有限公司,上海 201908;上海宝钢高新技术零部件有限公司,上海 201908;上海宝钢高新技术零部件有限公司,上海 201908;上海宝钢国际经济贸易有限公司,上海 200122【正文语种】中文【中图分类】TG162.79随着汽车工业的发展，世界各国对汽车的安全、绿色环保的要求越来越苛刻，采用超高强度钢是实现汽车减重、节能减排及提升安全的最经济性的选择。

然而，通常钢的成形性随着钢强度的提高而下降，以往通过将成形性和强化分为两个工艺步骤来实现。

热成形技术是实现超高强度汽车零件的一种全新工艺，通过热成形可以有效解决零件成形性和零件性能强化相矛盾的问题[1—2]。

稀土元素对轧制Mg–Zn–Zr_合金板材微观组织和力学性能的影响

128精密成形工程 2023年7月[9] HUANG Meng, XU Chao, FAN Guo-hua, et al. Role ofLayered Structure in Ductility Improvement of LayeredTi-Al Metal Composite[J]. Acta Materialia, 2018, 153: 235-249.[10] 王锟. 基于内聚力-GTN混合模型的钛-铝层状复合板损伤研究[D]. 洛阳: 河南科技大学, 2018: 43-46.WANG Kun. Study on Damage of Ti-Al Laminated Composite Plate Based on CZM-GTN Hybrid Model[D].Luoyang: Henan University of Science and Technology,2018: 43-46.[11] 郭照灿. 铜/铝双金属复层材料应变光学检测及力学性能研究[D]. 郑州: 郑州轻工业大学, 2022: 31-42.GUO Zhao-can. Optical Strain Detection and Mechani-cal Properties of Cu/Al Bimetallic Clad Materials[D].Zhengzhou: Zhengzhou University of Light Industry, 2022: 31-42.[12] 李莎, 贾燚, 刘欣阳, 等. 层状镁/铝复合板轧制工艺研究进展[J]. 精密成形工程, 2021, 13(6): 1-11.LI Sha, JIA Yi, LIU Xin-yang, et al. Research Progresson Rolling Process of Laminated Mg/Al Clad Plate[J].Journal of Netshape Forming Engineering, 2021, 13(6):1-11.[13] 高勃兴, 邹德坤, 谢红飙, 等. 铝/钢轧制复合有限元二次开发模拟与实验研究[J]. 精密成形工程, 2021, 13(6): 56-63.GAO Bo-xing, ZOU De-kun, XIE Hong-biao, et al.Simulation and Experimental Study on Finite Element Secondary Development of Aluminum/Steel Rolling Composite[J]. Journal of Netshape Forming Engineering, 2021, 13(6): 56-63.[14] LIU H S, ZHANG B, ZHANG G P. Enhanced Tough-ness and Fatigue Strength of Cold Roll Bonded Cu/Cu Laminated Composites with Mechanical Contrast[J].Scripta Materialia, 2011, 65(10): 891-894.[15] HUANG M, CHEN J S, WU H, et al. Strengthening andToughening of Layered Ti-Al Metal Composites by Controlling Local Strain Contribution[J]. IOP Confer-ence Series: Materials Science and Engineering, 2017, 219: 012028.[16] LI Z, LIN Y C, ZHANG L, et al. In-situ Investigation onTensile Properties of a Novel Ti/Al Composite Sheet[J].International Journal of Mechanical Sciences, 2022, 231: 107592.[17] HUANG C X, WANG Y F, MA X L, et al. InterfaceAffected Zone for Optimal Strength and Ductility inHeterogeneous Laminate[J]. Materials Today, 2018, 21(7): 713-719.[18] CHEN Wen-huan, HE Wei-jun, CHEN Ze-jun, et al.Extraordinary Room Temperature Tensile Ductility ofLaminated Ti/Al Composite: Roles of Anisotropy andStrain Rate Sensitivity[J]. International Journal of Plas-ticity, 2020, 133: 102806.[19] XING Bing-hui, HUANG Tao, XU Liu-jie, et al. Effectof Heat Treatment Process on the Microstructure of theInterface of Ti/Al Laminated Composite[J]. CompositeInterfaces, 2022, 29(7): 749-764.[20] 金属材料拉伸试验第1部分: 室温试验方法: GB/T228. 1—2010[S]. 2011.Metallic Materials-Tensile Testing-Part 1: Method ofTest at Room Temperature: GB/T 228.1—2010[S]. 2011.[21] 杨方方, 皇涛, 陈拂晓, 等. 异质金属层状复合板分层应力-应变关系研究[J]. 塑性工程学报, 2018, 25(1):187-191.YANG Fang-fang, HUANG Tao, CHEN Fu-xiao, et al.Research on Delamination Stress-Strain Relationship ofHeterostructure Laminated Composite Sheets[J]. Journalof Plasticity Engineering, 2018, 25(1): 187-191.[22] HUANG Tao, PEI Yan-bo, CHEN Fu-xiao, et al. ANovel Layered Finite Element Model for Predicting theDamage Behavior of Metal Laminated Composite[J].Composite Structures, 2023, 311: 116786.[23] TVERGAARD V, HUTCHINSON J W. The Relationbetween Crack Growth Resistance and Fracture ProcessParameters in Elastic-Plastic Solids[J]. Journal of theMechanics and Physics of Solids, 1992, 40(6): 1377-1397.[24] GURSON A L. Continuum Theory of Ductile Ruptureby Void Nucleation and Growth: Part I-Yield Criteriaand Flow Rules for Porous Ductile Media[J]. Journal ofEngineering Materials and Technology, 1977, 99(1): 2-15.[25] TVERGAARD V. Influence of Voids on Shear BandInstabilities under Plane Strain Conditions[J]. Interna-tional Journal of Fracture, 1981, 17(4): 389-407.[26] HUANG M, FAN G H, GENG L, et al. Revealing Ex-traordinary Tensile Plasticity in Layered Ti-Al MetalComposite[J]. Scientific Reports, 2016, 6(1): 1-10.责任编辑：蒋红晨第15卷第7期精密成形工程2023年7月JOURNAL OF NETSHAPE FORMING ENGINEERING129收稿日期：2023‒03‒22 Received ：2023-03-22基金项目：国家自然科学基金（52001188）Fund ：National Natural Science Foundation of China (52001188)作者简介：孙捷（1988—），男，博士，讲师，主要研究方向为镁合金塑性变形机理Biography ：SUN Jie (1988-), Male, Doctor, Research focus: plastic deformation mechanism of magnesium alloy. 引文格式：孙捷, 曲京儒, 阎玉芹, 等. 稀土元素对轧制Mg–Zn–Zr 合金板材微观组织和力学性能的影响[J]. 精密成形工程, 2023, 15(7): 129-135.SUN Jie, QU Jing-ru, YAN Yu-qin, et al. Effect of Rare Earth (RE) on Microstructure and Mechanical Property of Rolled 稀土元素对轧制Mg–Zn–Zr 合金板材微观组织和力学性能的影响孙捷，曲京儒，阎玉芹，赵彦华（山东建筑大学机电工程学院，济南 250101）摘要：目的为了使Mg–Zn–Zr 合金在热加工过后具有良好的力学性能及变形各向同性，在Mg–2Zn–0.5Zr 合金中添加不同含量的稀土元素，研究稀土元素对Mg–2Zn–0.5Zr 合金轧制后微观组织和力学性能的影响规律，以解决变形镁合金织构强、变形各向异性强的问题。

什么是Term Sheet

什么是Term Sheet嘟嘟【期刊名称】《数字财富》【年(卷),期】2001(000)002【摘要】<正> 前两年,韩先生和他的几个同学在海外注册了一间壳公司("A 公司"),并通过这间壳公司持有的中国子公司在国内开展业务。

最近为了引进资金,他们与一家海外投资公司("B公司")进行了频繁的接触。

经过初步评估,B公司认为A公司价值8000万美元,并有意向A公司投入2000万美元,同时成为A公司的优先股股东。

日前B公司给韩先生他们发来了"Term Sheet"。

尽管韩先生他们对B 公司的出价感到满意,但"Term Sheet"上满眼的陌生名词却令他们极为困惑。

"Term Sheet"直译为"合同条款清单"。

针对某项投资而言,"Term Sheet"可理解为投资方和被投资方就未来的投资交易所达成的【总页数】3页(P35-37)【作者】嘟嘟【作者单位】【正文语种】中文【中图分类】F830.49【相关文献】1.Sheet metal hardening curve determined by laminated sample and its adaptability to sheet forming processes [J],2.Formability enhancing of AZ31 magnesium alloy sheets by differential speed rollingFormability enhancing of AZ31 magnesium alloy sheets by differential speed rollingFormability enhancing of AZ31 magnesium alloy sheets by differential speed rollingFormability enhancing of AZ31 magnesium alloy sheets by differential speed rolling [J], XIA Wei-jun;CHEN Ji-hua;CHEN Ding3.Formability of the AMS 5596 Sheet in Comparison with EDDQ Steel Sheet [J], 无;4.The Effect of Variable Blank-Holder Forces on the Formability of Aluminum Alloy Sheets during Sheet Metal Forming [J], Sun Chengzhi Chen Guanlong Lin Zhongqin5.An Overlooked Term in Assessment of the Potential Sea-Level Rise froma Collapse of the West Antarctic Ice Sheet [J], Diandong Ren;Mervyn Lynch;Lance M. Leslie因版权原因，仅展示原文概要，查看原文内容请购买。

AMS ALLOY 上海慈东

|慈东钢种查询|产品中文产品类别慈东合金阿里巴巴慈东特钢阿里巴巴首页 >> 公司新闻 >> 产品知识AMS 相关高温耐蚀不锈钢及合金表AMS 相关高温耐蚀不锈钢及合金表USA AMS-Index-Nr.Metal SpecificationDelivery Forms UNS- Nummer AMS 5500 Alloy 30302 Laminated Sheet S30200 AMS 5501 Alloy 30304 Sheet, Strip, Foil S30400 AMS 5502 Alloy 5502 cancelled ./. AMS 5503 Alloy 51430 Sheet, Strip, Plate S43000 AMS 5504 Alloy 51410 Sheet, Strip, Plate S41000 AMS 5505 Alloy 51410 Mod Sheet, Strip, Plate S41000 AMS 5506 Alloy 51420 Sheet, Strip, Plate S42000 AMS 5507 Alloy 30316L Sheet, Strip, Plate S31603 AMS 5508 Alloy Greek Ascoloy Sheet, Strip, Plate S41800 AMS 5509 Alloy 5509 cancelled ./. AMS 5510 Alloy 30321 Sheet, Strip, Plate S32100 AMS 5511 Alloy 304L Sheet, Strip, Plate S30403 AMS 5512 Alloy 30347 Sheet, Strip, Plate S34700 AMS 5513 Alloy 30304 Sheet, Strip, Plate S30400 AMS 5514 Alloy 30305 Sheet, Strip, Plate S30500 AMS 5515 Alloy 30302 Sheet, Strip, Plate S30200 AMS 5516 Alloy 30302 Sheet, Strip, Plate S30200 AMS 5517 Alloy 30301 Sheet, Strip S30100 AMS 5518 Alloy 30301 Sheet, Strip S30100 AMS 5519 Alloy 30301 Sheet, StripS30100 AMS 5520 Alloy PH 15-7Mo Sheet, Strip, Foil, Plate S15700 AMS 5521 Alloy 30310S Sheet, Strip, Plate S31008 AMS 5522 Alloy 5522 cancelled ./. AMS 5523 Alloy 30309S Sheet, Strip, Plate S30908 AMS 5524 Alloy 30316 Sheet, Strip, Plate S31600 AMS 5525 Alloy A286 Sheet, Strip, Plate S66286 AMS 5526 Alloy 19-9DL Sheet, Strip, Plate K63198 AMS 5527 Alloy 5527 cancelled ./. AMS 5528 Alloy 17-7PH Sheet, Strip, Plate S17700 AMS 5529 Alloy 17-7PH Sheet, Strip S17700 AMS 5530 Alloy Hastelloy C Sheet, Strip, Plate N10002 AMS 5531 Alloy 5531 cancelled ./. AMS 5532 Alloy R30155 Sheet, Strip, Plate N-155 AMS 5533 Alloy 5533 cancelled ./. AMS 5534 Alloy 5534 cancelled ./. AMS 5535 Alloy 5535 cancelled ./. AMS 5536 Alloy Hastelloy X Sheet, Strip, Plate N06002 AMS 5537 Alloy L-605 Sheet, Strip, Foil, Plate R30605 AMS 5538 Alloy 5538 cancelled ./. AMS 5539 Alloy 5539 cancelled ./. AMS 5540 Alloy Inconel 600 Sheet, Strip, Plate N06600 AMS 5541 Alloy Inconel 722 Sheet, Strip N07722 AMS 5542 Alloy Sheet, Strip, Plate N07750 AMS 5543 Alloy 5543 cancelled ./. AMS 5544 Alloy Waspaloy Sheet, Strip, Plate N07001 AMS 5545 Alloy Rene 41 Sheet, Strip, Plate N07041 AMS 5546 Alloy AM-350 Sheet, Strip S35000 AMS 5547 Alloy AM-355 Sheet, Strip S35500 AMS 5548 Alloy AM-350 Sheet, Strip S35000 AMS 5549 Alloy AM-355 PlateS35500 AMS 5550 Alloy Inconel 702 Sheet, Strip, Plate N07702 AMS 5551 Alloy 5551 cancelled ./. AMS 5552 Alloy Incoloy 801 Sheet, Strip, Plate N08801 AMS 5553 Alloy Nickel 201 Sheet, Strip N02201 AMS 5554 Alloy AM-350Seamless TubingS35000AMS 5555 Alloy Nickel 205 Wire N02205 AMS 5556 Alloy 30347 Seamless or welded Tubing - Hydraulic Tubing S34700 AMS 5557 Alloy 30321 Seamless or welded Tubing - Hydraulic Tubing S32100 AMS 5558 Alloy 30347 Seamless or welded Tubing - Hydraulic Tubing S34700 AMS 5559 Alloy 30321 Seamless or welded Tubing - Hydraulic Tubing S32100 AMS 5560 Alloy 30304 Seamless or welded Tubing - Hydraulic Tubing S30400 AMS 5561 Alloy 21-6-9 Seamless or welded Tubing - Hydraulic Tubing S21900 AMS 5562 Alloy 21-6-9 Seamless or welded Tubing - Hydraulic Tubing S21904 AMS 5563 Alloy 30304 Seamless or welded Tubing - Hydraulic Tubing S30400 AMS 5564 Alloy 30304 Seamless or welded Tubing - Hydraulic Tubing S30400 AMS 5565 Alloy 30304 Seamless or welded Tubing - Hydraulic Tubing S30400 AMS 5566 Alloy 30304 Seamless or welded Tubing - Hydraulic Tubing S30400 AMS 5567 Alloy 30304 Seamless or welded Tubing - Hydraulic Tubing S30400 AMS 5568 Alloy 17-7PH Seamless or welded Tubing - Hydraulic Tubing S17700 AMS 5569 Alloy 30304I Seamless or welded Tubing - Hydraulic Tubing S30403 AMS 5570 Alloy 30321 Seamless or welded Tubing - Hydraulic Tubing S32100 AMS 5571 Alloy 30347 Seamless or welded Tubing - Hydraulic Tubing S34700 AMS 5572 Alloy 30310S Seamless or welded Tubing - Hydraulic Tubing S31008 AMS 5573 Alloy 30316 Seamless or welded Tubing - Hydraulic Tubing S31600 AMS 5574 Alloy 30309S Seamless or welded Tubing - Hydraulic Tubing S30908 AMS 5575 Alloy 30347 Seamless or welded Tubing - Hydraulic Tubing S34700 AMS 5576 Alloy 30321 Seamless or welded Tubing - Hydraulic Tubing S32100 AMS 5577 Alloy 30310 Seamless or welded Tubing - Hydraulic Tubing S31008 AMS 5578 Alloy Custom 455 Seamless or welded Tubing - Hydraulic Tubing S45500 AMS 5579 Alloy 19-9DL Seamless or welded Tubing - Hydraulic Tubing S63198 AMS 5580 Alloy Inconel 600 Seamless or welded Tubing - Hydraulic Tubing N06600 AMS 5581 Alloy Inconel 625 Seamless or welded Tubing - Hydraulic Tubing N06625 AMS 5582 Alloy Inconel X750 Seamless or welded Tubing - Hydraulic Tubing N07750 AMS 5583 Alloy Inconel X750 Seamless or welded Tubing - Hydraulic Tubing N07750 AMS 5584 Alloy 30316I Seamless or welded Tubing - Hydraulic Tubing S31603 AMS 5585 Alloy N-155 Seamless or welded Tubing - Hydraulic Tubing R30155 AMS 5586 Alloy Waspaloy Seamless or welded Tubing - Hydraulic Tubing N07001 AMS 5587 Alloy Hastelloy X Seamless or welded Tubing - Hydraulic Tubing N06002 AMS 5588 Alloy Hastelloy X Seamless or welded Tubing - Hydraulic Tubing N06002 AMS 5589 Alloy Inconel 718 Seamless or welded Tubing - Hydraulic Tubing N07718 AMS 5590 Alloy Inconel 718 Seamless or welded Tubing - Hydraulic Tubing N07718 AMS 5591 Alloy 51410 Seamless or welded Tubing - Hydraulic Tubing N41000 AMS 5592 Alloy RA 330 Sheet, Strip, Plate N08330 AMS 5593 Alloy RA 333 Sheet, Strip, Plate N06333 AMS 5594 Alloy 5594 cancelled ./.AMS 5595 Alloy 21-6-9 Sheet, Strip, Plate S21904 AMS 5596 Alloy Inconel 718 Sheet, Strip, Foil, Plate N07718 AMS 5597 Alloy Inconel 718 Sheet, Strip, Plate N07718 AMS 5598 Alloy Inconel X750 Sheet, Strip, Plate N07750 AMS 5599 Alloy Inconel 625 Sheet, Strip, Plate N06625 AMS 5600 Alloy 302 Laminated Sheet S30200 AMS 5601 Alloy 5601 cancelled ./.AMS 5602 Alloy 5602 cancelled ./.AMS 5603 Alloy PH14-8Mo Sheet, Strip - CANCELLED S14800 AMS 5604 Alloy 17-4PH Sheet, Strip, Plate S17400 AMS 5605 Alloy Inconel 706 Sheet, Strip, Plate N09706 AMS 5606 Alloy Inconel 706 Sheet, Strip, Plate N09706 AMS 5607 Alloy Hastelloy N Sheet, Strip, Plate N10003 AMS 5608 Alloy Haynes 188 Sheet, Strip, Plate R30188 AMS 5609 Alloy 51410 Mod Bars, Wire, Forgings, Tubing, Rings S41040 AMS 5610 Alloy 51416 / 51416Se Bars, Wire, Forgings S41623 AMS 5611 Alloy 51410 Mod Bars, Wire, Forgings, Tubing, Rings S4100I AMS 5612 Alloy 51410 Mod Bars, Wire, Forgings, Tubing, Rings S4100I AMS 5613 Alloy 51410 Bars, Wire, Forgings, Tubing, Rings S41000 AMS 5614 Alloy Bars, Wire, Forgings S41025 AMS 5615 Alloy 5615 cancelled ./.AMS 5616 Alloy Greek Ascoloy Bars, Wire, Forgings, Tubing, Rings S41800 AMS 5617 Alloy Cusiom 455 Bars, Wire, Forgings S45500 AMS 5618 Alloy 51440C Bars, Wire, Forgings S44004 AMS 5619 Alloy 5619 cancelled ./.AMS 5620 Alloy 51420F Bars, Wire, Forgings S42023 AMS 5621 Alloy 51420 Bars, Wire, Forgings S42000 AMS 5622 Alloy 17-4PH Bars, Wire, Forgings, Tubing, Rings, Forgings S17400 AMS 5623 Alloy Bars, Wire, Forgings, Tubing, Rings, Forgings K91456Alloy Bars, Wire, Forgings, Tubing, Rings, Forgings K91505AMS 5625 Alloy Bars, Wire, Forgings, Tubing, Rings, Forgings K91456AMS 5626 Alloy TI Bars, Wire, Forgings, Tubing, Rings, Forgings TI2001AMS 5627 Alloy 51430 Bars, Wire, Forgings, Tubing, Rings S43000AMS 5628 Alloy 51431 Bars, Wire, Forgings, Tubing S43100AMS 5629 Alloy PH 13-8Mo Bars, Forgings, Rings, Extrusions S13800AMS 5630 Alloy 51440C Bars, Wire, Forgings S44004AMS 5631 Alloy 51440A Bars, Forgings S44002AMS 5632 Alloy 51440FSe / 51440FBars, Wire, forgings S44020 / S44023 AMS 5633 Alloy 5633 cancelled ./.AMS 5634 Alloy 5634 cancelled ./.AMS 5635 Alloy 303Pb Bars, Wire, Forgings S30360AMS 5636 Alloy 30302 Bars, Wire S30200AMS 5637 Alloy 30302 Bars, Wire S30200AMS 5638 Alloy 303MA Bars, Forgings S30345AMS 5639 Alloy 30304 Bars, Wire, Forgings, Tubing, Rings S30400AMS 5640 Alloy 30303 Bars, Wire, Forgings S30300AMS 5640 Alloy 30303 Mod Bars, Wire, Forgings S30310AMS 5640 Alloy 30303 Se Bars, Wire, Forgings S30323AMS 5641 Alloy 30303 Se Bars, Wire, Forgings S30323AMS 5642 Alloy Bars, Wire, Forgings S34720 / S34723 AMS 5643 Alloy 17-4PH Bars, Wire, Forgings, Tubing, Rings S17400AMS 5644 Alloy 17-7PH Bars, Wire, Forgings, Tubing, Rings, Forgings S17700AMS 5645 Alloy 30321 Bars, Wire, Forgings, Tubing, Rings S32100AMS 5646 Alloy 30347 Bars, Wire, Forgings, Tubing, Rings S34700AMS 5647 Alloy 304L Bars, Wire, Forgings, Tubing, Rings S30403AMS 5648 Alloy 30316 Bars, Wire, Forgings, Tubing, Rings S31600AMS 5649 Alloy 316FM Bars, Wire, Forgings J92200AMS 5650 Alloy 30309S Bars, Wire, Forgings, Tubing, Rings S30908AMS 5651 Alloy 30310 Bars, Wire, Forgings, Tubing, Rings S31008AMS 5652 Alloy 30314 Bars, Wire, Forgings, Tubing, Rings S31400AMS 5653 Alloy 30316L Bars, Wire, Forgings, Tubing, Rings S31603AMS 5654 Alloy 30347 Bars, Wire, Forgings, Tubing, Rings S34700AMS 5655 Alloy 422 Bars, Wire, Forgings S42200AMS 5656 Alloy 21-6-9 Bars, Wire, Forgings, Extrusions, Rings S21904AMS 5657 Alloy PH15-7Mo Bars, Forgings S15700AMS 5658 Alloy 5658 cancelled S15500AMS 5659 Alloy 15-5PH Bars, Wire, Forgings, Extrusions, Rings S15500AMS 5660 Alloy Incoloy 901 Bars, Forgings N09901AMS 5661 Alloy Incoloy 901 Mod Bars, Forgings, Rings N09901AMS 5662 Alloy Inconel 718 Bars, Forgings, Rings N07718AMS 5663 Alloy Inconel 718 Bars, Forgings, Rings N07718AMS 5664 Alloy Inconel 718 Bars, Forgings, Rings N07718AMS 5665 Alloy Inconel 600 Bars, Forgings, Rings N06600AMS 5666 Alloy Inconel 625 Bars, Forgings, Extrusions, Rings N06625AMS 5667 Alloy Inconel X750 Bars, Forgings, Rings N07750AMS 5668 Alloy Inconel X750 Bars, Forgings, Rings N07750AMS 5669 Alloy Inconel X750 Bars, Forgings, Rings N07750AMS 5670 Alloy Inconel X750 Bars, Forgings, Rings N07750AMS 5671 Alloy Inconel X750 Bars, Forgings, Rings N07750AMS 5672 Alloy Custom 455 Bars, Forgings, Rings S45500AMS 5673 Alloy 5673 cancelled S17700AMS 5674 Alloy 30347 Wire S34700AMS 5675 Alloy FM92 Wire N07092AMS 5676 Alloy Nichrome V Wire N06003AMS 5677 Alloy Nichrome V cancelled N06003AMS 5678 Alloy 17-7PH Wire S17700AMS 5679 Alloy FM62 Wire N06062AMS 5680 Alloy SAE 30347 Wire S34781AMS 5681 Alloy 30347 Electrodes W34710AMS 5682 Alloy Nichrome V Rods, Wire ./.AMS 5683 Alloy 42 Rods, Wire ./.AMS 5684 Alloy Inconel WE 132 cancelled, Electrodes W86132AMS 5685 Alloy 30305 Rods, Wire S30500AMS 5686 Alloy 30305 Rods, Wire S30500AMS 5687 Alloy Inconel 600 Rods, Wire N06600AMS 5688 Alloy 30302 Rods, Wire S30200AMS 5689 Alloy 30321 Rods, Wire S32100AMS 5690 Alloy 30316 Rods, Wire S31600Alloy 30316 Rods, Wire W31610 AMS 5692 Alloy Rods, Wire S31680 AMS 5693 Alloy 30302 cancelled S30200 AMS 5694 Alloy Rods, Wire S31080 AMS 5695 Alloy 30310 cancelled W31010 AMS 5696 Alloy 30316 Rods, Wire S31683 AMS 5697 Alloy 30304 Rods, Wire S30400 AMS 5698 Alloy Inconel X750 Rods, Wire N07750 AMS 5699 Alloy Inconel X750 Rods, Wire N07750 AMS 5700 Alloy TPA Bars, Forgings, Rings S66009 AMS 5701 Alloy Inconel 706 Bars, Forgings, Rings N09706 AMS 5702 Alloy Inconel 706 Bars, Forgings, Rings N09706 AMS 5703 Alloy Inconel 706 Bars, Forgings, Rings N09706 AMS 5704 Alloy Waspaloy Bars, Forgings, RingsAMS 5705 Alloy CNS Bars, Forgings, Rings S63005 AMS 5706 Alloy Waspaloy Bars, Forgings, Rings N07001 AMS 5707 Alloy Waspaloy Bars, Forgings, Rings N07001 AMS 5708 Alloy Waspaloy Bars, Forgings, Rings N07001 AMS 5709 Alloy Waspaloy Bars, Forgings N07001 AMS 5710 Alloy XB Bars, Forgings S65006 AMS 5711 Alloy Hastelloy S Bars, Forgings, Rings N06635 AMS 5712 Alloy Rene 41 Bars, Forgings, Rings N07041 AMS 5713 Alloy Rene 41 Bars, Forgings, Rings N07041 AMS 5714 Alloy Inconel 722 Bars, Forgings, Rings N07722 AMS 5715 Alloy Inconel 601 Bars, Forgings, Rings N06601 AMS 5716 Alloy RA-330 Bars, Wire, Forgings, Rings N08330 AMS 5717 Alloy RA-333 Bars, Forgings, Rings N06333 AMS 5718 Alloy Jethele M-152 Bars, Forgings, Tubing, Rings S64152 AMS 5719 Alloy Jethele M-152 Bars, Wire, Forgings, Rings, Extrusions S64152 AMS 5720 Alloy 5720 cancelled ./.AMS 5721 Alloy 5721 cancelled ./.AMS 5722 Alloy 5722 cancelled ./.AMS 5723 Alloy 5723 cancelled ./.AMS 5724 Alloy 5724 cancelled ./.AMS 5725 Alloy 5725 cancelled ./.AMS 5726 Alloy A286 Bars, Wire S66286 AMS 5727 Alloy 5727 cancelled ./.AMS 5728 Alloy 5728 cancelled ./.AMS 5729 Alloy 5729 cancelled ./.AMS 5730 Alloy 5730 cancelled ./.AMS 5731 Alloy A286 Bars, Wire, Forgings, Tubing, Rings S66286 AMS 5732 Alloy A286 Bars, Wire, Forgings, Tubing, Rings S66286 AMS 5733 Alloy Discaloy Bars, Forgings S66220 AMS 5734 Alloy A286 Bars, Wire, Forgings, Tubing S66286 AMS 5735 Alloy Bars, Forgings, Tubing, Rings K66286 AMS 5736 Alloy Bars, Forgings, Tubing, Rings K66286 AMS 5737 Alloy A286 Bars, Wire, Forgings, Tubing S66286 AMS 5738 Alloy 30303F Bars, Wire, Forgings, Tubing S30323 AMS 5739 Alloy Almar 362 Bars, Forgings, Rings S36200 AMS 5740 Alloy Almar 362 Bars, Forgings, Rings S36200 AMS 5741 Alloy 5741 cancelled ./.AMS 5742 Alloy 5742 cancelled ./.AMS 5743 Alloy AM-355 Bars, Forgings S35500 AMS 5744 Alloy AM-355 Bars, Forgings S35500 AMS 5745 Alloy AM-350 Bars, Forgings S35000 AMS 5746 Alloy D-979 Bars, Forgings N09979 AMS 5747 Alloy Inconel X750 Bars, Forgings, Rings N07750 AMS 5748 Alloy 5748 cancelled ./.AMS 5749 Alloy BG-42 Bars, Wire, Forgings, Tubing S42700 AMS 5750 Alloy Hastelloy C Bars, Forgings, Rings N10002 AMS 5751 Alloy Udimet-500 Bars, Forgings, Rings N07500 AMS 5752 Alloy 5752 cancelled ./.AMS 5753 Alloy 5753 cancelled ./.AMS 5754 Alloy Hastelloy X Bars, Forgings, Rings N06002 AMS 5755 Alloy Hastelloy W Bars, Forgings, Rings N10004 AMS 5756 Alloy 5756 cancelled ./.AMS 5757 Alloy 5757 cancelled ./.AMS 5758 Alloy MP-35N Bars, Wire, Forgings, Tubing R30035 AMS 5759 Alloy L-605 Bars, Forgings, Rings R30605AMS 5760 Alloy 5760 cancelled ./.AMS 5761 Alloy 5761 cancelled ./.AMS 5762 Alloy 203EZ Bars, Wire, Forgings S20300AMS 5763 Alloy Custom 450 Bars, Forgings, Tubing, Rings S45000AMS 5764 Alloy 22-13-5 Bars, Wire, Forgings, Extrusions, Rings S20910AMS 5765 Alloy 5765 cancelled ./.AMS 5766 Alloy Incoloy 800 Bars, Forgings N08800AMS 5767 Alloy 5767 cancelled ./.AMS 5768 Alloy N-155 Bars, Wire, Forgings, Rings R30155AMS 5769 Alloy N-155 Bars, Forgings, Rings R30155AMS 5770 Alloy 5770 cancelled ./.AMS 5771 Alloy Hastelloy N Bars, Forgings, Rings N10003AMS 5772 Alloy Haynes 188 Bars, Forgings, Rings R30188AMS 5773 Alloy Custom 450 Bars, Wire, Forgings, Tubing, Rings S45000AMS 5774 Alloy AM-350 Bars, Wire, Forgings, 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S15789AMS 5813 Alloy 5813 cancelled ./.AMS 5814 Alloy Wire R30918 AMS 5817 Alloy Greek Ascoloy Wire S41800AMS 5818 Alloy WireAMS 5819 Alloy Wire R30021AMS 5821 Alloy 51410 Mod Wire S41081AMS 5822 Alloy Wire S41780AMS 5822 AAlloy Wire S41780 AMS 5823 Alloy Wire S41780 AMS 5824 Alloy 17-7-PH Wire S17780AMS 5825 Alloy 17-4 PH Wire S17480AMS 5826 Alloy 15-5PH Wire S15500AMS 5827 Alloy 5827 cancelled ./.AMS 5828 Alloy Waspaloy Wire N07001AMS 5829 Alloy Nimonic 90 Wire N07090AMS 5830 Alloy Inocoloy 901 Wire N09901AMS 5831 Alloy HS-556 Wire R30556AMS 5832 Alloy Inconel 718 Wire N07718AMS 5833 Alloy Elgiloy Wire R30003AMS 5834 Alloy Elgiloy Wire R30003AMS 5835 Alloy WireAMS 5836 Alloy Inconel FM-82 Wire N06082AMS 5837 Alloy Inconel 625 Wire N06625AMS 5838 Alloy Hastelloy S Wire N06635AMS 5839 Alloy Haynes 230W Wire N06231AMS 5840 Alloy PH13-8Mo Wire S13889AMS 5841 Alloy MP159 Bars, Forgings R30159AMS 5842 Alloy MP159 Bars, Forgings R30159AMS 5843 Alloy MP159 Bars, Forgings R30159AMS 5844 Alloy 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Influence of out-of-plane compression stress on limit strains in sheet metals

ORIGINAL RESEARCHInfluence of out-of-plane compression stress on limit strains in sheet metalsMorteza Nurcheshmeh &Daniel E.GreenReceived:10December 2010/Accepted:18March 2011/Published online:5May 2011#Springer-Verlag France 2011Abstract The prediction of the forming limits of sheet metals typically assumes plane stress conditions that are really only valid for open die stamping or processes with negligible out-of-plane stresses.In fact,many industrial sheet metal forming processes lead to significant compres-sive stresses at the sheet surface,and therefore the effects of the through-thickness stress on the formability of sheet metals cannot be ignored.Moreover,predictions of forming limit curves (FLC)that assume plane stress conditions may not be valid when the forming process involves non-negligible out-of-plane stresses.For this reason a new model was developed to predict FLC for general,three-dimensional stress states.Marciniak and Kuczynski (Int J Mech Sci 9:609-620,1967)first proposed an analytical method to predict the FLC in 1967,known as the MK method,and this approach has been used for decades to accurately predict FLC for plane stress sheet forming applications.In this work,the conventional MK analysis was extended to include the through-thickness principal stress component (σ3),and its effect on the formability of different grades of sheet metal was investigated in terms of the ratio of the third to the first principal stress components (b ¼s 3s 1=).The FLC was predicted for plane stress conditions (β=0)as well as cases with different compres-sive through-thickness stress values (β≠0)in order to study the influence of βon the FLC in three-dimensional stress conditions.An analysis was also carried out to determine how the sensitivity of the FLC prediction to the through-thickness stress component changes with variations in the strain hardening coefficient,in the strain rate sensitivity,in plastic anisotropy,in grain size and in sheet thickness.It was found that the out-of-plane stress always has an effect on the position of the FLC in principal strain space.However,the analysis also showed that among the factors considered in this paper,the strain hardening coefficient has the most significant effect on the dependency of FLC to the through-thickness stress,while the strain rate sensitivity coefficient has the least influence on this sensitivity.IntroductionThe poor correlation between the common “cupping ”test and the actual performance of sheet metal in industrial forming operations led researchers to look at some more fundamental parameters.A significant breakthrough came in 1963,when Keeler and Backofen [2]reported that during sheet stretching,localized necking required a critical combination of major and minor strains (along two perpendicular directions in the plane of the sheet).Subsequently,this concept was extended by Goodwin [3]to sheet drawing and the resulting curve is known as the Keeler-Goodwin curve or the forming limit curve (FLC).In other words,Keeler developed the right side of the FLC (i.e.,positive minor strain),and Goodwin extended the forming limit curve to include negative minor strains.In order to predict the FLC,Marciniak and Kuczynski [1]proposed that the inhomogeneity of the sheet material could be modeled by a geometric defect in the sheet.In their study,an imperfection in the form of a shallow grooveM.Nurcheshmeh (*):D.E.Green Department of Mechanical,Automotive and Materials Engineering,University of Windsor,401Sunset Avenue,Windsor,Ontario N9B 3P4,Canada e-mail:nurches@uwindsor.caURL:http://www.uwindsor.ca/engineering/Int J Mater Form (2012)5:213–226DOI 10.1007/s12289-011-1044-9Keywords Forming limit curve .Out-of-plane stress .Formability .MK analysis .Sheet metalwas applied to specimens stretched in equibiaxial tension. The severity of the imperfection was quantified by the ratio of the thickness in the groove to the nominal thickness of the sheet.In general,no reductions in the forming limit would be seen when the value of the imperfection factor is between0.99and1.00.In this model,the initial inhomo-geneity of the material develops continuously with plastic deformation until a localized neck eventually appears.In1970,Azrin and Backofen[4]subjected a large number of materials to in-plane stretching.They discovered that an imperfection factor of about0.97or less was required to obtain agreement between the MK analysis and experimental FLC data.Accordingly,even though the MK method provided a simple predictive model,there was inconsistency between its predictions and experimental data.Similar trends were also observed by Sowerby and Duncan[5]as well as by Marciniak et al.[6].In addition, Sowerby and Duncan[5]also reported that limit strains predicted with the MK method showed a considerable dependence on material anisotropy.Ghosh[7]found that material strain rate sensitivity is important during post-uniform deformation.The additional hardening due to strain rate sensitivity plays a significant role in increasing the forming limits by delaying strain localization inside the neck.The physical soundness and the simplicity of the MK analysis has no doubt been the reason this method has been the most popular theoretical approach for FLC calculation, and it has been used by many researchers,even in recent years:for instance Butuc et al.[8]in2006,Yoshida et al.[9]in2007and Nurcheshmeh and Green[10]in2010.The prediction of the FLC of sheet metals traditionally assumes plane stress loading conditions and the effect of the normal stress is usually neglected.Therefore FLC predictions are only strictly valid for open die and free forming processes.However,many metal forming pro-cesses lead to the development of non-negligible normal stresses in the sheet when it is formed over a die radius. Through-thickness stresses become even more significant in hydroforming processes,where a pressurized fluid compresses a sheet or a tube against the surface of the die.In many hydroforming applications,the pressure of the forming fluid can generate such high contact pressures that the through-thickness stress exceeds the in-plane stresses.The existence of a significant through-thickness compressive stress creates a hydrostatic stress state that has the potential to increase the formability of the sheet and therefore requires consideration in the prediction of the FLC.Very few sheet formability studies have taken into account the effect of the normal stress and further research is required in this area.Gotoh et al.[11]presented an analytical expression that predicts an increase in the plane-strain forming limit in strain space due to the presence of through-thickness compressive stresses.They demon-strated theoretically that an out-of-plane stress(even as small as one tenth of the yield stress)can raise the forming limit strain and thus can be effectively used to delay the onset of fracture in press forming.Smith et al.[12]developed a new sheet metal formability model that takes into account the through-thickness normal stress for materials that exhibit planar isotropy.These authors’model predicts a greater increase in formability due to compressive stresses than that predicted by Gotoh’s model.They also examined the influence of the strain hardening coefficient(n value)on the sensitivity of the FLC to the normal stress.Finally,Banabic and Soare[13]used the MK analysis to study the influence of fluid pressure normal to the sheet surface on the forming limits of thin,orthotropic sheets. Their model was used to predict the FLC of AA3104-H19 aluminum alloy subject to different fluid pressures ranging from0(plane stress condition)to200MPa.They showed that the formability of this aluminum alloy improves with the application of a fluid pressure,especially on the right side of the forming limit diagram.Experimental data was available in the plane stress condition which was predicted satisfactory and used to calibrate their model.In the present paper,a three-dimensional stress state was implemented in a modified version of the MK model to predict FLC with different through-thickness stress values. The imperfection factor was related to the surface rough-ness and grain size of the sheet and was updated throughout the deformation of the sheet.The imperfection band was oriented perpendicular to the first principal stress,and its rotation was also considered as the sheet was plastically deformed.This modified MK model was validated in plane-stress conditions with experimental FLC data obtained for AISI-1012steel[14]and it was also compared with other theoretical results obtained by the present authors[10].The validation of the model for cases that involved through-thickness stresses was done with pub-lished experimental FLC data for AA6011aluminum[15] and STKM-11A steel[16]sheets.The sensitivity of the predicted FLC to the applied out-of-plane stress component was also analyzed as a function of variations in different material properties and the results of this sensitivity analysis will be discussed.Theoretical approachMarciniak and Kuczynski[1,6]presented a theoretical framework for prediction of FLC that is commonly known as the MK method,which has been shown to predict FLCs with reasonable accuracy.This approach is based on thefact that inhomogeneities are unavoidable in actual sheet materials,and it is assumed that this inherent material inhomogeneity can be modeled as a geometric imperfection in the form of a narrow band(Fig.1)with a slightly different thickness than the rest of the sheet.Although this approach was originally proposed for plane stress con-ditions,the current work includes the third stress compo-nent in the MK model and is shown asσ3in Fig.1.Figure1schematically represents a shallow groove on sheet surface,which effectively divides it into two separate regions:region(a)with nominal thickness,and region(b)with the reduced thickness in the groove.The initial imperfection factor of the groove,f0,is defined as the thickness ratio between the two regions as follows:f0¼t b0að1Þwhere t denotes the sheet thickness and subscript‘0’denotes the initial state.The thickness difference between these two regions is critical element in the MK theory because the predicted limiting strains are very sensitive to the initial value of the imperfection factor.In most studies,this coefficient is simply assumed to have a fixed value close to1.0and that can be adjusted so that the predicted FLC will better fit the experimental data. However,it has been proposed[10]that a more realistic approach would be to relate the initial thickness differ-ence between the two regions to the surface roughness of the sheet.Indeed,research carried out by Stachowicz[17] shows that surface roughness changes with deformation and these changes depend upon initial surface roughness, grain size,and effective plastic strain.By relating the thickness difference between regions(a)and(b)to the surface roughness of the sheet metal,the imperfection factor not only takes on a value that has physical meaning but also the option of adjusting this value so that the predicted FLC can better fit experimental data is elimi-nated.Stachowicz’s assumption was adopted in this work and the imperfection factor was assumed to change with the deformation of the sheet according to the following relationship:f0¼t a0À2R Z0þCd0:50"b eÂÃt a0ð2aÞf¼t a0À2R Z0þCd0:50"b eÂÃt a0exp"b3À"a3ÀÁð2bÞwhere R Z0is the surface roughness before deformation,C is a material constant,"b e is the effective strain in region(b), and d0is the material’s initial grain size.Additional details on the calculation of the imperfection factor are provided in the authors’previous work[10].In general,the imperfection band is randomly oriented and its orientation can be determined by the angleθbetween the groove axis and the direction of the second principal stress (Fig.1).When plastic deformation begins,this angle will slowly start to change as the groove rotates with respect to the loading axes,and its orientation can affect the limiting strains. In order to obtain FLC predictions with good accuracy,the variations in the groove orientation should therefore be considered in the calculation of the forming limit strains by updating its value at each increment throughout the plastic deformation.This rotation of the imperfection band during deformation was well researched by Sing and Rao[18]and they proposed an empirical formula in which the orientation varies as a function of the true plastic strain increments in region(a)of the sheet as follows:tan qþd qðÞ¼tan qðÞ1þd"a11þd"a2ð3Þwhere d"a1and d"a2are the major and minor principal strains in the nominal area of the sheet,respectively.A constitutive equation was derived in which the yield function can be expressed in the following general form for isotropic hardening:f¼32S ij:N:S ij1=2Às eð4Þwhere,S is the deviatoric stress tensor and N is a tensor that describes the anisotropy of the sheet material in terms of the anisotropic constants in Hill’s1948yield function[19].With consideration of the third principal stress compo-nent,the three-dimensional plastic potential function was implemented in the MK analysis:2h¼s x2þFþHðÞs y2þFþGðÞs2zÀ2H s x s yÀ2F s z s yÀ2G s z s x¼f2ð5Þwhere the anisotropic coefficients F,G and H can be calculated from the yield stresses in the principaldirections. Fig.1Thickness imperfection in the MK modelStrain hardening is described with the power hardening including strain rate sensitivity effect as follows:s e ¼k "•e m"e þ"0ðÞnð6Þwhere ε0is a uniform prestrain applied to the sheet,m is thestrain-rate sensitivity coefficient,n is the strain-hardening coefficient,σe and εe are the effective stress and strain,respectively.The associated flow rule was employed to calculate plastic strain increments as follows:d "ij ¼d l Âgrad ðh Þ¼d l Â@h @s ijð7Þwhere d 1is the plastic multiplier and h is the plasticpotential function.There are two main assumptions in the MK analysis.The first one is the geometric compatibility equation expressed as the equality of the tangential plastic strain components inside and outside the imperfection band,d "a tt¼d "b ttð8Þand the second assumption is the equilibrium of the normaland shear forces across the imperfection,i.e.:F a nn ¼F b nnð9a ÞF a nt ¼F bnt ð9b Þwhere subscripts n and t denote the normal and tangential directions of the groove,respectively,and F is the force per unit width,i.e.:F a nn ¼s a nn ta ð10a ÞF b nn ¼s b nn t b ð10b ÞF a nt ¼s a nt ta ð10c ÞF b nt ¼s b nt t bð10d ÞBy combining Eqs.1,6and 10a ,10b the following relation is obtained:s a nn s e!s b nns e!0¼f "0þ"b e ÂÃn Â"b e &m "0þ"a e ÂÃn Â"a e &m0ð11a ÞSince the strain rate is defined as "e &¼d "e dt =,it follows that:s a nn s a e !s b nn s b e!0¼f "0þ"b e ÂÃ="0þ"a e ÂÃÀÁn Âd "b e =d "a eÂÃmð11b ÞFinally,the stress transformation rule leads to the expressions:s a nn ¼s a x cos 2q ðÞþs a y sin 2q ðÞð12a Þs a nt¼Às a x Às aysin q ðÞcos q ðÞ¼s a x a À1ðÞsin q ðÞcos q ðÞ½ð12b Þwhere αis the ratio of the second true principal stresscomponent (σ2)to the first true principal stress component (σ1)in the nominal area which indicates the stress path.Expressions similar to Eqs.12a and 12b can be written for region (b),and using Eqs.9,10,and 12we obtain:s b nt s nn ¼s a nts nn¼a À1ðÞsin q ðÞcos q ðÞcos q ðÞþa sin 2q ðÞð13ÞWith consideration of the consistency condition,theplastic potential function and the strain transformation rule:d "ae s eF þH ðÞÂa a ÀF b a ÀH ½ s a x cos 2q þ1ÀG b a ÀH a a ðÞs a x sin 2q ÈÉ¼s b xd "b es b eF þH ðÞÂa b ÀF b b ÀH ÂÃcos 2q þ1ÀG b b ÀH a b ÀÁsin 2q ÈÉð14Þwhere βis the ratio of the third true stress component to thefirst true stress component,such that:b ¼s 3s 1¼s z s x==ð15ÞBy combining Eqs.11,13,and 14,the final governingequation was analytically determined as a function of the ratio of the effective plastic strain inside and outside the imperfection band h ¼"b e "a e .This final differential equa-tion indicates the evolution of the effective plastic strain ratio ηas the sheet is deformed under a three-dimensional loading condition.The plastic deformation of the sheet begins as strain increments are imposed along a linear strain path (i.e.for a constant value of r ¼"2"1=)in the nominal region,and the stress components are calculated from the strain state in the nominal area.Then the strains and stresses in the imperfection region are calculated from the strains and stresses in the nominal area by using the governing equations described above.During the analysis,it is assumed that the normal stress applied on the surface ofthe sheet or tube is identical for both region(a)and region(b)of the MK model.But since the thickness in region(b)is less than that in the rest of the sheet,the strain rateincreases faster in region(b)than in region(a).Moreover,the difference in strain rate between the two regions willintensify as the deformation progresses,and eventually thestrains will localize in the imperfection region.It isgenerally assumed that plastic instability occurs when theeffective plastic strain in the imperfection region reachesten times that in nominal area("b e¼10"a e).Once the onset of necking takes place,the in-plane plastic strain compo-nents in the nominal area("a1and"a2)identify a point on theFLC for the specified strain pathρ.In order to generate theentire FLC,the value of the strain ratioρis modified andthe procedure is repeated for each new strain path.The FLCis thus determined from the limiting strain data obtained forstrain paths that vary in incrementsΔρ=0.05from uniaxialtension(ρ=−0.5)to equibiaxial tension(ρ=1.0). Experimental validation of the modified MK modelThe theoretical MK analysis model presented in the previous section was implemented into a numerical code.This proposed model was then used to predict the FLC of actual sheet and tube materials,both with and without applied normal stresses,in order to validate the numerical code. Description of materialsThe materials that were considered for the validation of the proposed MK model are a low carbon steel(AISI-1012) [14],AA6011aluminum alloy[15],and STKM-11A steel [16](the designation of this last steel grade follows the Japanese standard and it is equivalent to an MT1010steel in the ASTM standard).The mechanical properties of these materials are listed in Table1.It is also worth noting that in these publications,AISI-1012refers to a flat stock sheet metal,whereas AA6011and STKM-11A refer to thin walled tubes.Equation2a was used to calculate the initial imperfection factor value in the MK analysis.It was found that f0=0.995 for AISI-1012steel,f0=0.997for AA6011aluminum,and f0=0.991for STKM-11A steel.Validation of the proposed MK modelIn order to validate the three-dimensional FLC model described in the previous section,theoretical FLCs were calculated in both plane stress and three-dimensional stress conditions and the predicted FLCs were compared with published experimental data[14–16].The new model was verified first under plane stress conditions,in the absence of through-thickness stresses (β=0).Theoretical FLC were compared with the experimen-tal FLC of as-received AISI-1012sheet steel[14]which were obtained by carrying out stretch forming tests using rectangular and notched blanks of various widths with different conditions of lubrication to achieve a range of strain statesÀ0:5r¼"2"11:0=.Each blank was electro-etched with a 3.0mm diameter circle grid and formed over a hemispherical punch until the onset of local necking.The major and minor strains were measured directly from the deformed grids using a profile projector. The FLC predicted with the proposed MK model was also compared with the FLC predicted by a different MK analysis code developed previously by the same authors for purely plane stress conditions[10].The predicted and experimental FLCs for this grade of steel are shown in Fig.2.Figure2shows good agreement between the theoretical and experimental FLCs obtained under plane stress con-ditions,and the developed model predicts the FLC for this steel with acceptable accuracy.Furthermore,it can be seen that the FLC predicted under plane stress conditions with the new three-dimensional model is essentially identical to the FLC predicted with the previous two-dimensional analysis code[10].The proposed MK analysis model was also verified for more general loading conditions where the out-of-plane stress component is non-negligible(β≠0).This further validation of the three-dimensional MK model was carried out by predicting the FLC of AA6011aluminum tubes that were hydroformed with up to15-MPa internal pressure (which corresponds toσ3≈7.5MPa).Hwang et al.[15] prepared200-mm long tube specimens with a1.86-mm wall thickness,and a51.9-mm outer diameter.The tube specimens were annealed at410°C for2h and then a grid of5-mm-diameter circles with a spacing of1-mm was electrochemically etched onto the surface of undeformed tubes for the purpose of strain measurement.Tubes were pressurized in a bulge test apparatus without axial feeding to generate positive minor strains.Other tubes were also pressurized in a hydroforming test machine with axial feeding to generate strain paths with negative minor strains. After the tubes were deformed,the circle grids in the vicinity of the burst were measured by a three-dimensional digital image processing system and the major and minorTable1Mechanical propertiesMaterial K(MPa)n m R(Normal)t0(mm)AISI-1012[14]2380.350.015 1.21 2.5 AA6011[15]254.90.265–0.574 1.86 STKM-11A[16]14500.14– 2.14 1.4strains were determined.The limiting strain data from these tests was used to construct the left side of the FLC of these aluminum tubes.The comparison of the predicted and experimental FLCs is shown in Fig.3.It can be seen from Fig.3that there is good agreement between the experimental data and the predicted FLC on the left side of the diagram.This may seem surprising considering that the analysis was carried out using Hill ’s 1948yield criterion.Indeed,it is well known that Hill ’s quadratic yield function is not suitable for predicting the biaxial behaviour of aluminum alloys and more recent,non-quadratic yield functions have been shown to be much more appropriate [20].However,it can be seen that the experimental FLC data in Fig.3corresponds with defor-mation modes between plane strain and uniaxial tension,and for such deformation modes the quadratic yield function is capable of predicting reasonably accurate results.Non-quadratic yield functions typically lead to improved predictions of the forming behaviour of alumi-num alloys for deformations in biaxial tension,because they are better able to represent the shape of the yield locus between plane strain and balanced biaxial tension:this corresponds with the right side of the FLC for which no experimental data is available.No doubt the predictions of FLC in the region of plane strain would be improved with the use of a non-quadratic yield function.The proposed model was also validated with another set of experimental limiting strain data for STKM-11A steel presented by Kim et al.[16].These authors determined the experimental FLC by hydroforming straight tubes with bothMinor StrainFig.2Comparison of predicted and experimental FLCs of AISI-1012steel sheet in-plane stress condition [14]Minor StrainM a j o r S t r a i nFig.3Comparison of predicted and experimental FLCs ofAA6011aluminum sheets under 15MPa internal pressure [15]an axial end-feed force and 56-MPa internal pressure (leading to σ3≈28MPa).A constant ratio of high internal pressure and relatively low axial force was applied with an end displacement rate of 2.33-mm/s using a PC-based controller.During these experiments,tubes were pressur-ized until they burst,and the average burst pressure was 56MPa,with the split occurring parallel with the tube axis and positioned toward the middle of the tube.Strain measurements were taken as near to the fractured edge as possible in order to determine limit strains.Figure 4shows a comparison of predicted and experimental FLC for negative minor strains.It can be seen in Fig.4that the FLC predicted by the proposed MK analysis lies slightly above the experimental FLC for this grade of steel.This discrepancy between the theoretical and the experimental FLC data is likely due tothe fact that experimental strains were not actually measured in local necks since these tubes were allowed to burst,but they were measured in the uniformly deformed material right next to the fractured edge of burst tubes.Therefore these experimental strain data represent a conservative estimation of the actual FLC.Limiting strain data was not available for the right hand side of the diagram because Kim et al.[16]were only able to apply a compressive axial force to the ends of the tubes,whereas a tensile axial force is required to obtain positive minor strains [21].It is also worth pointing out that the experimental FLC data [14–16]used to validate the current MK model were obtained using the well-known circle grid analysis tech-nique.This technique relies on the measurement of deformed grids on the surface of the specimens as well asMinor StrainM a j o r S t r a i nFig.4Comparison of predicted and experimental FLCs of STKM-11A steel sheet under 56MPa internal pressure [16]Minor StrainM a j o r S t r a i nFig.5FLC of AISI-1012sheet steel predicted as a function of the applied normal stressthe somewhat subjective interpretation about whether necking has begun or not in a specific grid location.This technique is therefore dependent on the experimentalist ’s experience and the accuracy of the strain measurements,and therefore it inevitably leads to some variability in the results.According to the author ’s experience,the experimental error that can be expected in FLC strain data obtained with the circle grid technique is estimated to be within ±2.5%strain.More advanced techniques are now being used to determine the forming limits of sheet materials with greater repeatability and reproducibility.For instance,digital image correlation is used to measure the strain field across the entire specimen gauge area and numerical interpolation methods are then used to determine the strains at the onset of necking [22–26].These techniques are very powerful as they can determine limiting strains even for very high strength materials that tend to fracture without necking.However,although there is some experimental error in the published experimental FLC data [14–16],the comparisons between the predicted and experimental FLC (Figs.2,3and 4)nevertheless show that the proposed three-dimensional MK model provides a good prediction of the FLC,whether the through-thickness stress component is significant or not.Influence of the through-thickness stress on the FLC The primary purpose of this work is to study the effect of the through-thickness stress component on the forming limit curve.In this section,the sensitivity of the FLC to the out-of-plane stress component will be studied by applying different levels of through-thickness stress to the surface of AISI-1012steel sheets.The FLC was predicted for a normal stress ranging from σ3=0(plane stress condition)to σ3=35MPa.The theoretical results are presented in Fig.5.It can be seen from Fig.5that the FLC is quite sensitive to the normal stress:indeed,the entire FLC is observed to shift up the vertical axis when the applied normal stress increases.The formability of this sheet steel is seen to improve with a normal stress as low as 10MPa.Furthermore,it is apparent from Fig.5that the increase in formability is not proportional to the increase in normal stress:indeed,the rate of increase in formability also increases with the normal stress.Influence of mechanical properties on the sensitivity of FLC to out-of-plane stressesIn the previous section it was shown (Fig.5)that the FLC of AISI-1012sheet steel is dependent on the magnitude ofMinor StrainM a j o r S t r a i nFig.6FLC of a sheet material that differs from AISI-1012only by its strain hardening coefficient (n=0.70),predicted as a function of the applied normalstress510152025303540Through Thickness Stress (MPa)% i n c r e a s e i n F L D oFig.7Increase in FLC0as a function of the applied normal stress for two sheet steels that differ only by their strain hardening coefficient (n=0.35and n=0.70)。

201防锈油中英文版.xls

201 防锈油201 anti-rust oil本产品采用深度精制基础油，加入黄酸钡、黄酸钠、羊毛脂等多种添加产品执行国家标准。

The product is using depth of refined base oils and with baryta yellow, POTA IUM AMYLIC X, lanoline and other various additive products, and is executing GB standard.典型数据typical data分析项目analyze project 分析结果Analysisresults试验方法test method运动粘度kenematicviscosity28.53 GB17265闪点（开口）flash point 170 GB173536凝点solidifying point -10 GB173535酸值acid number（加添加剂前）（before additiveproducts）0.025 GB1712581机械杂质mechanicalimpurity0.02 GB17511水份moisture 痕迹vestige GB17260水溶性酸或碱watersoluble acid or alkali中性neutral人汗置换性sweatsubstitution 45号钢片45# steel 72 hours合格pass盐水浸渍试验salt waterimmersion test45#steel 100hours1 GB175096湿热试验damp heattest 45# steel 96 hours 1用途：application本产品适用于机械设备，机械零部件等金属表面，有良好的防锈作用。

This product is applicable to mechanical equipment, machinery parts and other metal surface, have good anti-rust effect.注意事项：considerations使用及储运过程中要防止异物污染以及水混入，不要与其他油品混用，防止性能变差。

中英文外文翻译---笔记本上盖外壳的镁合金薄板冲压模具设计

Die design for stamping a notebook case with magnesium alloy sheetsContent SummaryIn the present study，the stamping processfor manufacturing anotebook top cover case with LZ91 magnesium–lithium alloy sheet at roomtemperature was examined using both the experimental approach and the finite element analysis. A four-operation stamping process was developed to eliminate both the fracture and wrinkle defects occurred in the stamping process of the top cover case. In order to validate the finite element analysis，an actua four-operation stamping process was conducted with the use of 0.6mm thick LZ91 sheetas the blank. A good agreement in the thickness distribution at various locations between the experimental data and the finite element results confirmed confirmed the accuracy and efficiency of the ementanalysis.The super or for mability of LZ91 sheet at room temperature was also demonstrated in the present study by successful manufacturing of the notebook topcover case. The proposed four operation process lend sit selftoan efficient approach to form the hinge in the notebook with less number of operational procedures than that required in the current practice. It also confirms that the notebook cover cases can be produced with LZ91 magnesium alloy sheet by the stamping process. It provides an alternative to the electronics industry in the application of magnesium alloys. Keywords: Notebook case；LZ91 magnesium–lithium alloy sheet；stamping；Multi-operation；Formability1. IntroductionDue to It slight weight and good performance in EMI resistance, magnesium alloy has been widely used for structural components in the electronics industry, such as cellular phones and notebook cases. Although the prevailing manufacturing process of magnesium alloy products has been die casting，the st- amping of magnesium all sheet has drawn interests from industry because of its competitive productivity and performance in the effective production of thin-walled structural components.As for stamping process，AZ31 magne siumalloy (aluminum 3%, zinc 1%) sheet has been commonly used for the for ming process at the present time,even though it needs to be formed at elevated temperature due to its hexagonal closed packed (HCP) crystal structureRecently,the magnesium–lithium(LZ)alloy has also been successfully deve- loped to improve the formability of magnesium alloy at room temperature. The ductility of magnesium alloy can be improved with the addition of lit hium that develops the formation of body centered-cubic (BCC) crystal structure (Takuda et al., 1999a,b; Drozd et al,2004).In the present study, the stamping process of a notebook top cover case with the use of LZ sheet was examined. The forming of the two hinges in the top cover of a notebook, as shown in Fig.1(a and b),is the most difficult operation in the stamping process due to the small distance between the flanges and the small corner radii at the flanges, as displayed in Fig. 1(c). This geometri complexity was caused by a dramatic change in the corner radius when the flange of get stooclo set the notebook,which would easily cause fracture defect around the flange of hinge and requirea multi-operation stamping process to overcome this problem.In the present study, the formability of LZ magnesium alloy sheets was invest- igated and an optimum multi-operation stamping process was developed to reduce the number of operation all proced using both the experiment approach and the finite element analysis.Fig.1–Flange of hinges at notebook top cover case.(a) Hinge, (b) top cover case and (c) flanges of hinge.2. Mechanical properties of magnesiumcontent of lithium increases. It is also observ from Fig. 2(a) that the curves of LZ91 sheet at room temperature and AZ31 sheet at 200,C are close to each other. LZ101 sheet at room temperature exhibit seven better ductility than LZ91 and AZ31 do at 200,C. Since the cost of lithium is very expensive, LZ91 sheet, instead of LZ101 sheet, can be considered as a suitable LZ magnesium alloy sheet to render favorable formability at room temperature. For this reason ,the present study adopted LZ91 sheet as the blank for the notebook top cover case and attempted to examine the formability of LZ91 at room temperature. In order to determine if the fracture would occur in the finite element analysis, the forming limit diagram for the 0.6mm thick LZ91 sheet was also established as shown in Fig. 2(b).alloy sheets The tensile test swereper formed for magnesium–lithiumalloy sheets of LZ61 (lithium 6%, zinc 1%), LZ91, and LZ101 at room temperature to compare their mechanical properties to those of AZ31 sheets at elevated temperatures. Fig. 2(a) shows the stress–strain relations of LZ sheets at room temperature and those of AZ31 sheets at both room temperature and 200?C. It is noted that the stress–strain curve tends to be lower.Fig. 2 – Mechanical properties of magnesium alloy.(a) The stress–strain relations of magnesium alloy; (b) forming limit diagram (FLD) of LZ91 sheet.3. The finite element modelThe tooling geometries were constructed by a CAD software, PRO/E, and were converted into the finite element mesh ,as shown in Fig. 3(a), using the software DELTAMESH. The tooling was treated as rigid bodies, and the four-node shell element was adopted to construct the mesh for blank. The material lproper ties and forming limitd iagram sobtained from the experiments were used in the finite element simulations. The other simulation parameters used in the initial run were: punch velocity of 5mm/s, blank-holder force of 3kN, and Coulomb friction coefficient of 0.1. The finite element software PAM STAMP was employed to perform the analysis, and the simulations were performed on a desktop PC.A finite element model was first constructed to examine the oneoperation forming process of the hinge. Due to symmetry, only one half of the top covercase was simulated, as showninFig.3(a).The simulation result, as show ninFig.3(b),indicates that fracture occurs at the corners of flanges, and the minimum thickness is less than 0.35mm. It implies that the fracture problem is very serious and may not be solved just by enlarging the corner radii at the flanges. The finite element simulation swere performed to study the parameters .That affect the occurrence off racture. Several approaches were proposed to avoid the fracture as well.Fig. 3 – The finite element simulations. (a) Finite element mesh and (b) fracture at the corners.4. Multi-operation stamping process designIn order to avoid the occurrence of fracture, a multi-operation stamping process is required. In the current industrial practice, itusually take satle ast tenoperational procedures to form the top cover case using the magnesium alloy sheet. In thepresent study, attempts were made to reduce the number of operational procedures. Several approaches were proposed to avoid the fracture, and the four-operation stamping process had demonstrated itself as a feasible solution to the fracture problem. To limit the length of this paper, only the two operation and the four-operation stamping processes were depicted in the following.4.1 Two-operation stamping processThe first operation in the two-operation stamp in process was side wall forming as shown in Fig.4(a),and the second one was the forming off lange ofhing epresented in Fig.4(b),the height of the flange of hinge being 5mm .Fig.4(c)shows the thickness distribution obtained from the finite element simulation. The minimum thickness of the deformed sheet was 0.41mm and the strains were all above the forming limit diagram. It means the fractured effect could be avoided. Inaddition, the height of the flange conformed to the target goal to be achieved. How- ever, this process produced a critical defect of wrinkling, as shown in Fig. 4(d), on the flange of hinge, which induces a problem in the subsequent trimming operation. Hence, even though the two-operation stamping process solved the fracture problem at the corner of the bottom and the flange of hinge, a better forming process is still expected to solve the wrinkling of flange of hinge.Fig. 4 – Two-operation stamping process.(a) Formation of sidewalls, (b) formation of hinges, (c) thickness distribution and (d) wrinkle.4.2. Four-operation stamping processThe four-operation forming process proposed in the present study starts with the forming of three side wall sand the flange of the hinge with a generous corner radius, as shown in Fig.5(a).Since the side wall close to the flange was open and the corner radius was larger than the desired ones, theflange was successfully formed without fracture. Such process success-fully avoided the difficulty of forming two geometric features simultaneously, but increased the material flow of the blank sheet. The next step was to trim the blank outside the side walls, and to calibrate the corner radius of 4mm to the desired value of 2.5mm. The hinge was thus formed, as shown in Fig. 5(b). The third step was to fold the open side, so that the sidewall could be completed around its periphery, as shown in Fig. 5(c). The effect of trimming the extra sheet outside the sidewalls in the second step on the third step was studied. When the extra sheet was not trimmed, the thickness at the corner was 0.381mm, as shown in Fig. 5(d). The thickness of Table Comparison of thickness measured ABCD Experiment 0.42mm 0.44mm 0.49mm 0.53mm Simulation 0.423mm 0.448mm 0.508mm 0.532mm Error 0.71% 1.79% 3.54% 0.38% the corner increased to 0.473mm, as shown in Fig. 5(e), if the trimming was implemented in the second step. The excessive material producedby the folding process in the third step was then trimmed off according to the parts design. The last step was the striking process that is applied to calibrate all the corner radii to the designed values. The minimum thickness at the corner of the final product was 0.42mm,and all the strains were above the forming limit diagram. It is to be noted that Fig. 5(a–c) only shows the formation of one hinge. The same design concept was then extended to the stamping process of the complete top cover case.5. Experimental validationIn order to validate the finiteel ement analysis,an actualfour operation stamping process was conducted with the use of 0.6mm thick LZ91 sheet as the blank. The blank dimension and the tooling geometries were designed according to the finite element simulation results. A sound product without fracture and wrinkle was then manufactured, as shown in Fig. 6(a). To further validate the finite element analysis quantitatively, the thickness at the corners around the hinge of the sound product, as shown in Fig. 6(b), were measured and compared with those obtained from the finite element simulations, as listed in Table 1. It is seen in Table 1 that the experimental data and the finiteelement results were consistent. The four-operation process design based on the finite element analysis was then confirmed by the experimental data.Fig. 6 – The sound product. (a) Without fracture and wrinkle and (b) locations of thickness measured.Concluding remarksThe press forming of magnesium alloy sheets was studied in the present study using the experimental approach and the finite element analysis. The formability of both AZ31 and LZ sheets was examined first. The research results in dicated th a the LZ91 sheet has favorable formability at room temperature, which is similar to that of AZ31 sheet at the forming temper- ature of 200C.The superior formability of LZ91 sheet at room tempera Ture was also demonstrated in the present study by successful manufacturing of the notebook top cover case. The proposed four-operation process lends itself to an efficient approach to form the hinge in the notebook with fewer operational procedures than that required in the current practice. It also confirms that the notebook cover cases can be produced with LZ91 magne siumalloy LZ91sheet by the stamping process. It provides an alternative to the electronics industry in the application of magnesium alloys. Acknowledg ments The authors would like to thank the National Science Council of the Republic of China for financially supporting this research under the Project No. NSC-95-2622-E-002-019-CC3, which made this research possible. They would also like to thank ESI, France for the help in running the PAM STAMP program.References[1] Chen. F.K.Huang.T.B.Chang. C.K.2003. Deep drawing of square cups with magnesium alloy AZ31sheets. Int. J. Mach. Tools[2] Manuf. 43.1553–1559.Drozd.Z..Trojanova′ .Z, Ku′ dela.S.2004. Deformation of behavior ofMg–Li–Al alloy. J. Mater. Compd. 378. 192–195.[3]Takuda.H.Yoshii.T. Hatta, N.1999a. Finite-element analysis of the formability of a based alloy AZ31sheet. J.[4] Mater. Process. Technol. 89/90. 135–140.Takuda.H. Kikuchi.S.[5]Tsukada.T.Kubota.K.Hatta.N.1999b.Effect of strain rate on deformation behavior of a Mg–8.5Li–1Zn alloy sheet at room temperature. Mater. Sci. Eng. 271, 251–256.笔记本上盖外壳的镁合金薄板冲压模具设计内容提要在本研究中，在室温下分别用实验方法和有限元分析对笔记本上盖的lz91镁合金薄板冲压工艺制造情况进行检查。

S-121-01铝件钝化外观标准

铝件钝化外观规范S-121-011 范围索斯科公司生产的所有钝化铝件的要求.2 目的此份文件概述了索斯科和相关等级系统,种类,检查标准和可接受标准. 适用于采购件和生产件.所有生产部门和供应商都须参照此份说明;3 有争议时的优先处理方法当此份说明和其它的采购信息来源发生矛盾时,按以下方法解决:3.1 除非有正确的文件支持,如偏差,样品等3.2 有合适的索斯科的工程图纸3.3 以此份外观说明为准4. 表面分级的定义A 级: 在最终用户安装完成后总是可以看到的高装饰性的区域。

B级: 在最终用户安装完成后总是可以看到的非装饰性的区域。

C级: 在最终用户安装完成后经常可以看到的区域。

D级: 在最终用户安装完成后很少或不能够看到的区域。

5 设计者/工程的职责5.1 在图纸上指出表面等级。

5.2 对于此份说明中未定义出的另外的外观要求须在图纸上标识出来。

5.3 图纸无外观要求的应按等级B检验。

5.4 对于单个产品的表面可用幻影线分出不同部分的表面等级。

6.0 观察条件6.1 所有检查需在正常的光源下目视进行,6.2 目视不能在放大镜下进行。

如果最终用途未知或多样，则目视应与表面成45度角。

6.3 等级A检查时，应反复调整保证表面在光源下的反射最强，便于观察。

6.4 等级B，C，D检查时，不需要反复调整。

7.0不良品的定义:7.1 刮伤:-表面浅的凹槽，包括凹槽形成的线;7.2凹痕: 表面上小的碗形坑;7.3 压痕: 由碰撞或挤压形成的孔或洞;7.4 表面异物: 表面明显来自外面的块状或片状物；7.5 毛刺：产品边缘的粗糙边脊或突起;7.6 边缘突起: 边部多余的镀层；7.7起泡：圆型缺失或不良，由于基材和金属沉淀之间空心而生成；7.8 露底: 产品某个区域缺少镀层；7.9 生锈/锈蚀：基材或镀层本身被氧化；7.10剥离：电镀层于基材分离或部分分开；7.11小孔/气孔：延伸至基材的镀层横断面上的一条不连续的圆形缝隙；7.12水纹：白色半透明或云状的区域，有一定的形状或形态；7.13赃点：掉色或被污染的点；7.14角落白/角落灼伤：常见于产品外角的发白现象；7.15 漏镍: 由于铬镀层没有盖住下面的镍层而引起的黄斑；7.16 指印：正常制作过程中留下的指纹痕迹；7.17 模印：印痕统一的刮花，由形成产品的模具引起；7.18 缩水: 产品表面的凹陷7.19 飞边：熔融塑胶结合形成部件处；7.20 夹水线: 塑胶件的变色或光泽改变，通常发生在水口/厚度突变处/通道阻隔附近；7.21 顶白拉白：产品合模线周围的多余材料；7.22 焦痕：产品表面的褐色斑点或飞纹。

sus201不锈钢化学成分欧洲标准

sus201不锈钢化学成分欧洲标准一、概述在工业生产和日常生活中，不锈钢已经成为了一个不可或缺的材料。

不锈钢因其耐腐蚀、美观、易清洁等特点，在家居用品、厨具、建筑材料等各个领域得到了广泛的应用。

而sus201作为不锈钢的一种，其具体的化学成分和物理性能对于使用者来说尤为重要。

在本文中，我们将对sus201不锈钢的化学成分及其符合的欧洲标准进行详细介绍。

二、sus201不锈钢的化学成分sus201是一种含氮奥氏体型不锈钢，其主要成分包括：1. 铬（Cr）：17.00-19.002. 镍（Ni）：3.50-5.503. 锰（Mn）：5.50-7.504. 硅（Si）：1.00以下5. 碳（C）：0.15以下6. 磷（P）：0.060以下7. 硫（S）：0.030以下8. 氮（N）：0.25以下9. 铜（Cu）：0.80以下三、sus201化学成分的作用1. 铬：铬是不锈钢的主要合金元素，对不锈钢的耐腐蚀性起到关键作用。

适当的铬含量能够形成一层致密的氧化膜，阻止金属内部继续被腐蚀。

2. 镍：镍的主要作用是提高不锈钢的耐腐蚀性，增强其抗氧化性，并且能够提高不锈钢的塑性。

3. 锰：锰的加入可以改善不锈钢的耐蚀性和强度，同时也能够提高不锈钢的塑性和焊接性。

4. 硅：硅是不锈钢中的元素之一，其加入可以提高不锈钢的硬度和强度。

5. 碳：碳的含量对不锈钢的抗拉强度有影响。

6. 磷和硫：磷和硫能够影响不锈钢的加工性，同时也会影响其耐蚀性。

7. 氮：氮的加入能够提高不锈钢的强度和耐磨性。

8. 铜：铜的加入可以改善不锈钢的耐蚀性和耐磨性。

四、sus201不锈钢符合的欧洲标准据欧洲标准，sus201不锈钢的化学成分应符合以下要求：1. 铬含量应在16.0-18.0之间。

2. 镍含量应在3.5-5.5之间。

3. 锰含量应在5.5-7.5之间。

4. 硅含量应在1.0以下。

5. 碳含量应在0.15以下。

6. 磷含量应在0.060以下。

【轻量化】孙黎-诺贝丽斯新一代高强车用铝板及应用

1 Next Generation High Strength Aluminum Alloys for Automotive ApplicationsLi SunNovelis China R&D1 Novelis is the #1 producerof flat-rolled aluminum sheet with a global market share of~50%40We have morethan 40 years’experience in theautomotivemarket134We have morethan 134automotive specificpatents225Novelis aluminumcan be found inmore than 225vehicle modelsaround the worldNOVELIS AUTOMOTIVE TECHNOLOGY LEADERSHIPNovelis Advanz™ family of alloys is designed for the next generation of vehiclesEuropeAsiaNovelis is the world’s largest supplier of aluminum sheet to the automotive industry with R&T, operations and commercial teams strategically located across North America, Europe and Asia.North AmericaEuropeAsiaNorth America$300 Million, 200KMT/yr. automotive finishing lines to open in 2020$180 Million,100KMT/yr automotive finishing line to open in 2020Automotive Strategy5Our product portfolio consists of the latest advances in product development, including:Series alloys have excellent strength-to-weight ratio, formability properties, and full recycling compatibility5xxxSeries alloys are versatile,heat-treatable, highly formable,weldable6xxxSeries alloys with high-energy management characteristics for increased crash performance and superior in-servicestrength7xxxOur proprietary process that simultaneously casts multiple alloylayers into a single aluminum rollingingotFusionOUR PRODUCT PORTFOLIONovelis’ Portfolio of AlloysStandard ProductsNovelis Global ProductsNovelis New Product LaunchesExteriorInner/StructureStrength Crash575451826111RC5754RC5182Advanz™ e170Advanz™ c300Advanz™ e600Advanz™ s118Advanz™ s600Advanz™ e200Advanz™ s200Advanz™ s615Advanz™ s650Advanz™ s701Advanz™ Fusion e2008Comparing Select Novelis Alloys with Steel Grades for Automotive Applications010********20406080100120140160180200E l o n g a t i o n t oF r a c t u r e (%)Specific Yield Strength (MPa /cm 3/g)DP TRIP CP FB HSLA BH IFDQDuctibor MS57545182HF6Xs615s650s701s702Comm ercialR&DSteel GradesAutomotive Strategy: Advanced Alloys9Automotive Strategy: Key Enablers1011Automotive Strategy: Key Enablers –Recycling。

基于AUTOFORM的汽车前围横梁连接板的数值分析与实验研究

基于AUTOFORM的汽车前围横梁连接板的数值分析与实验研究徐迎强;韩永志;崔礼春【摘要】目的分析汽车前围横梁连接板的冲压成形过程.方法以板料成形非线性分析软件AUTOFORM为平台,对汽车前围横梁连接板的冲压成形过程进行CAE分析.根据模拟结果(成形极限图、材料流动分布及材料变薄率),对拉延型面及工艺参数进行了优化.结果所得零件材料最大减薄率为14.6％,在B340LA(t=1.0 mm)材质减薄率安全范围内(16.9％),零件型面球化处角部无暗伤及拉裂,翻边处材料流动均匀,无开裂风险,成形结果得到大大改善.结论 CAE仿真能够预测零件成形过程中存在的缺陷,优化工艺参数,指导模具设计工作.最后将优化结果用于指导实际生产,得到了符合质量要求的零件.【期刊名称】《精密成形工程》【年(卷),期】2014(006)003【总页数】5页(P15-19)【关键词】汽车前围横梁连接板;数值模拟;工艺补充;拉延成形【作者】徐迎强;韩永志;崔礼春【作者单位】安徽江淮汽车股份有限公司,合肥230601;安徽江淮汽车股份有限公司,合肥230601;安徽江淮汽车股份有限公司,合肥230601【正文语种】中文【中图分类】TG386车身覆盖件尺寸较大、结构复杂，为空间曲面形状，用简单的数学解析很难表达，其加工过程涉及几何、材料非线性和复杂的摩擦状态等问题。

由于影响因素较多，因此无法精确控制材料的流动，较难找到变形规律，出现的质量问题较多[1—6]。

传统意义上的试模法造成大量资源浪费，已无法应对如今产品更新换代的短周期的要求，板料成形CAE使工艺设计人员可以在投产准备阶段，预估零件成形中可能出现的缺陷，如回弹、起皱及破裂等，并优化工艺参数，验证并指导后期的模具设计工作[7—13]。

前围横梁连接板的成形工艺，由拉延(DR)、修边(TR)、翻边(FL)及整形(RST)等4道工序完成，其中DR是关键，它决定了TR，FL和RST等工序的内容及成形状态。