Research progress of high-entropy amorphous materials and their additive manufacturing technology
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摘要: 高熵非晶合金具有独特的物理、化学和力学性能以及更好的热稳定性，因而其制备技术成为国内外重要的研究热点之一. 然而利用传统技术制备高熵非晶材料时会产生晶粒粗大及材料浪费等缺点，难以满足工艺生产需要. 而增材制造技术的精准制造和快速冷却等特点可以解决这一问题，制备出各项性能优越的高熵非晶合金. 简要介绍了高熵非晶材料的研究体系和常用制造方法，着重阐述了高熵非晶材料的断裂强度、耐腐蚀性和热稳定性的研究，对增材制造技术的工艺特征和优势，以及利用增材制造技术制备高熵非晶合金的科学难点作出了总结. 结果表明，利用增材制造技术有利于获得致密均匀的高熵非晶材料，但对于非晶相形成的解释仅限于高熵合金4大效应.最后阐述了近年来利用常用的两种增材制造手段制造高熵非晶合金的研究，并对增材制造技术制备高熵非晶材料的发展趋势提出了展望.Abstract: High-entropy amorphous alloys (HEAAs) exhibit unique physical, chemical and mechanical properties as well as better thermal stability. Thus, its fabrication technology has become one of the important research hotspots at home and abroad. However, high-entropy amorphous materials manufactured by traditional technology had defects such as coarse crystal grains and material waste, which was difficult to meet the needs of processing production. The precise manufacturing and rapid cooling of additive manufacturing technology could solve the problems, and produce high entropy amorphous alloys with superior properties. This review research briefly introduced the research system and common preparation methods of high-entropy amorphous materials. It mainly focused on the research about fracture strength, corrosion resistance and thermal stability of high-entropy amorphous materials. The process features and advantages of additive manufacturing technology, and the scientific difficulties for applying this technology to fabricate high-entropy amorphous alloys were summarized. The results showed that additive manufacturing technology contributed to high-entropy amorphous materials with dense and uniform microstructures, while the explanation for the formation of amorphous phases was limited to the four effects of high-entropy alloys, Finally, a discussion with two additive manufacturing methods commonly used in the fabrication of high-entropy amorphous materials in recent years was made. Furthermore, the prospects for the development trend of fabricating high-entropy amorphous materials by additive manufacturing technology were put forward.
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Keywords:
- high-entropy alloy /
- amorphous phase /
- additive manufacturing /
- grain refinement
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0. 序言
X80管线钢(规定总延伸强度Rt0.5 > 555 MPa)是目前世界范围内大输量、高压油气输送管道的首选钢种，自1985年德国Ruhr Gas /Megal Ⅱ 项目中首次成功铺设了ϕ1 118 mm × 3.2 km × 13.6 mm (壁厚)的试验段以来[1]，国内外已建成多条X80管线钢长距离油气输送管道[2-3]. 基于油气输送管道高输送压力、超大输量、大管径、高强度级别的发展趋势，管道建设和运行对管线钢的焊接效率和质量、安全性和经济性等提出了更高的要求，开发和应用新型高效、高质量焊接方法已成为必然趋势. 激光电弧复合焊技术(hybrid laser arc welding, HLAW) 是近年来发展的一种先进焊接方法，具有大熔深、快焊速、低变形、高质量等优势[4-6]，已成为管道施工焊接技术领域的极具良好应用前景的施工焊接方法. 但由于HLAW方法常采用较大的焊接速度，焊件冷却较快，可能导致高强钢焊接接头产生淬硬组织和较高的残余应力水平. 因此，分析X80管线钢中厚板激光电弧复合焊接头的显微组织构成及残余应力分布，对研究和控制X80级及以上强度级别管线钢HLAW接头的显微组织、焊接过程中的冶金反应、焊后残余应力和变形，具有重要的理论指导意义.
目前为止，国内外对高强钢激光电弧复合焊的数值模拟研究，主要针对采用较低激光功率焊接的温度场、流场的分析，而对较大板厚、较大激光功率的复合焊接接头的残余应力和变形的研究则十分缺乏. Kim等人[7]研究了激光/CO2电弧的热输入比值变化对7 mm厚SM 490钢激光电弧复合焊平板堆焊件的残余应力和变形的影响，发现激光和电弧热输入相同时角变形最小，总热输入不变时两者能量比对焊接残余应力分布几乎没有影响. Zhang等人[8]研究了10 mm厚激光-MAG (metal active gas, MAG)焊对接接头的残余应力分布和产生机理，认为峰值温度和冷却速度对残余应力的影响较大，马氏体相变可以降低残余应力水平. 严春妍等人[9]研究了焊接速度对12 mm厚S355钢激光-MIG (metal inert gas, MIG)单道焊接头残余应力分布的影响规律，发现纵向残余应力、横向残余应力峰值均随着焊接速度的增加而升高. 对激光电弧复合焊接残余应力和变形的研究多数为板厚较小、采用单道焊的情况，研究结论无法适用于采用多道多层焊厚大结构件的残余应力分析.
文中对18.4 mm厚的X80管线钢板进行激光-MIG多道焊试验，分析复合焊接头中不同区域的显微组织特征，并采用SYSWELD有限元软件对多道激光-MIG复合焊焊接过程的温度场和接头的残余应力进行计算，研究不同激光功率下激光-MIG复合焊接头残余应力分布的规律和特征.
1. 试验方法
激光-MIG复合焊接试验采用YSL-10000-S2型光纤激光器和福尼斯TPS-5000焊接电源，保护气体氩气的流量为25 L/min. 采用尺寸为180 mm × 60 mm × 18.4 mm的X80管线钢板进行平板对接，焊丝为林肯Merit JM-80低合金钢焊丝，母材和焊丝的化学成分如表1所示. 采用8 mm钝边的30°V形坡口，坡口间隙为0.5 mm，激光在前引导电弧. 环境温度和焊件初始温度均为20 ℃，层间温度为150 ℃. 试验采用的焊接工艺参数如表2所示.
表 1 X80钢和焊丝的主要化学成分(质量分数，%)Table 1. Main chemical compositions of X80 piepeline steel and filler wire材料 C Si Mn P S Cr Ni Mo X80 0.030 0.240 1.710 0.008 0.002 0.022 0.250 0.190 JM-80 0.081 0.670 1.560 0.018 0.008 0.330 0.011 0.008 表 2 X80管线钢复合焊工艺参数Table 2. HLAW welding parameters of X80 pipeline steel焊接方法 试样编号 激光功率
PL/kW光丝间距
d/mm焊接电流
I/A电弧电压
U/V焊接速度
v/(m·min−1)离焦量
△f0/mm打底焊 S1,S2,S3,S4,S5 9.0 2.0 190 23 1.2 −1 填充焊
盖面焊S1 2.0 2.0 230 24 0.5 0 S2 2.5 2.0 230 24 0.5 0 S3 3.0 2.0 230 24 0.5 0 S4 3.5 2.0 230 24 0.5 0 S5 4.0 2.0 230 24 0.5 0 采用上海点应光学体视显微镜测量焊缝的熔宽和熔深，采用JSM-6360LV型扫描电子显微镜对复合焊接头的显微组织进行观察. 采用USB-4718型温度采集卡测量焊件上表面典型位置的热循环曲线，采用ASMB6-60型应力应变采集器测试焊件上表面的残余应力.
2. 数值模拟模型
2.1 几何模型
基于试验测得所有焊道的形貌参数，采用SolidWorks软件建立5个激光功率下复合焊试样的实体模型，并进行有限元网格划分. 基于查阅的数据采用插值法来获得X80管线钢的热物理性能和力学性能数据[10]，对SYSWELD软件的材料数据库二次开发后进行多道激光-MIG复合焊的温度场和应力应变场的计算. 图1为复合焊接试样S2对应的有限元网格模型，可以看出共3道焊缝. 其余4个功率下的焊件也均获得3道焊缝，只是填充焊缝和盖面焊缝的形貌尺寸(熔深、熔池长度、熔池宽度等)存在差别.
2.2 热源模型
试验采用高8 mm钝边的30°V形坡口，为获得和实际焊接熔池形貌、尺寸接近的模拟结果，选用双椭球热源模型+三维锥体热源模型进行组合[9]，并基于焊道体积的变化对施加到焊道单元上的热源模型参数进行调整. 计算时采用生死单元模拟多道焊过程，将式(1) ~ 式(3)所示的热源模型在时间上离散，施加在各焊道对应的单元上进行计算.
电弧热源部分采用双椭球热源的热流密度函数，如式(1) ~ 式(2)所示[9]，即
$$ {q_{\text{f}}}(x,y,z) = \frac{{6\sqrt 3 Q{f_1}\sin \beta \sin \gamma }}{{{\text{π}} {a_1}bc\cos \theta \sqrt {\text{π}} }} \exp \left[ { - \frac{{3{x^2}}}{{{{({a_1}/\sin \beta )}^2}}} - \frac{{3{y^2}}}{{{{(b/\sin \gamma )}^2}}} - \frac{{3{z^2}}}{{{{(c\cos \theta )}^2}}}} \right] $$ (1) $$ {q_{\text{r}}}(x,y,z) = \frac{{6\sqrt 3 Q{f_2}\sin \beta \sin \gamma }}{{{\text{π}} {a_2}bc\cos \theta \sqrt {\text{π}} }} \exp \left[ { - \frac{{3{x^2}}}{{{{({a_2}/\sin \beta )}^2}}} - \frac{{3{y^2}}}{{{{(b/\sin \gamma )}^2}}} - \frac{{3{z^2}}}{{{{(c\cos \theta )}^2}}}} \right] $$ (2) 式中：qf和qr为热源前部和后部的热流密度；f1和f2为双椭球热源前、后两部分的能量分布，且f1+f2 = 2；Q为双椭球热源对应的有效热输入；a1，a2，b，c为双椭球热源模型尺寸参数；β，γ，θ分别为x，y，z 3个方向和电弧主轴的夹角.
激光热源部分对应的三维锥体热源模型的热流密度函数如式(3)所示[9]，即
$$ {q_{{\text{laser}}}}(x,y,z) = {q_0}\exp \left[ - \frac{{{x^2} + {y^2}}}{{{r_0}^2(z)}}\right] $$ (3) $$ {r_0}(z) = {r{{\rm{_e}}}} + \frac{{{r{{\rm{_i}}}} - {r{{\rm{_e}}}}}}{{{z{{\rm{_i}}}} - {z{{\rm{_e}}}}}}(z - {z{{\rm{_e}}}}) $$ (4) 式中：q0为激光热源的有效功率；re和ri分别为锥体热源上、下表面的热源半径；ze和zi分别为锥体热源上、下表面在高度方向z轴的坐标.
3. 结果与讨论
3.1 显微组织
对5个接头试样的显微组织进行分析发现，打底焊缝因工艺参数相同，显微组织非常接近，焊缝和粗晶热影响区(coarse grained heat affected zone, CGHAZ)中形成较多的条状贝氏体(lath bainite, LB)；不同激光功率PL下填充和盖面焊对应的焊缝、CGHAZ和细晶热影响区(fine-grained heat affected zone, FGHAZ)的显微组织类型基本相同，仅各相的比例有一定差别. 为便于分析，以S2试样(PL = 2.5 kW)为例，对其焊接接头盖面层的显微组织进行分析，如图2所示.
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图 2 盖面层的焊缝区和热影响区显微组织(S2试样)Figure 2. Microstructures in weld metal and HAZ of cap pass (specimen S2). (a) arc zone weld; (b) arc zone CGHAZ; (c) arc zone FGHAZ; (d) laser zone weld; (e) laser zone CGHAZ; (f) laser zone FGHAZ从图2可以看出，电弧区、激光区焊缝的显微组织均含有较多的针状铁素体(acicular ferrite, AF)、先共析铁素体(proeutectoid ferrite, PF)、侧板条铁素体(ferrite side plate, FSP)和粒状贝氏体(granular bainite, GB)，但激光区焊缝出现少量的LB；电弧区、激光区CGHAZ显微组织均主要由LB+GB组成，电弧区CGHAZ晶粒更加粗大、并出现了少量的AF；电弧区、激光区FGHAZ晶粒较小，显微组织均由GB+准多边形铁素体(quasi-polygonal ferrite, QPF)组成. 随着PL增加，焊缝中AF数量增加、GB和LB数量减少，CGHAZ中GB，AF数量增加，FGHAZ中QPF数量略有增加.
由于5组试验中填充、盖面焊缝所用的MIG焊焊接电流、电弧电压、焊接速度均相同，因此只研究盖面焊缝熔深Hc和激光作用部位HAZ宽度WL随PL的变化规律，如图3所示. 随着PL从2.0 kW增加到4.0 kW，Hc和WL分别增加了27.7%和31.9%，说明激光功率增加，引起盖面焊缝下半部分热输入增加，对增加盖面焊缝的熔深和HAZ宽度的作用十分明显.
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图 3 激光功率对盖面焊缝形貌参数的影响Figure 3. Influence of laser power on cap weld geometry parameters3.2 温度场计算结果
图4为试样S2的熔池形貌模拟结果和试验结果的对照图，可以看出二者形状比较吻合. 试样S2的每层焊缝填充过程的温度场计算结果如图5所示. 以试样S2与试样S5为例，复合焊接试样表面典型位置(P1点位于盖面焊缝表面中心线；P2，P3点分别位于焊件表面距熔合线1.5，2.5 mm处)的热循环曲线如图6所示.
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图 4 试验与模拟所得焊缝截面比较Figure 4. Comparison of experimental obtained and numerical predicted sectional morphology![]()
图 5 激光-MIG复合焊接温度场(试样S2)Figure 5. Temperature field of laser-MIG welding of specimen S2. (a) root layer; (b) filler layer; (c) cap layer![]()
图 6 不同PL下典型位置的热循环曲线Figure 6. Thermal cycles of representative points with different PL. (a) P1; (b) P2 and P3图5所示的计算结果表明，当激光功率PL为2.5 kW时，打底焊的焊接温度场的峰值温度为2 211.26 ℃，填充焊和盖面焊的焊接温度场峰值温度略有升高，分别达到2 356.15和2 425.36 ℃. 当PL从2.5 kW升高至4.0 kW时，3个典型位置热循环曲线的峰值温度均有所升高. P1点热循环曲线的峰值温度从1 988.63 ℃上升至2 404.82 ℃，P2点热循环曲线的峰值温度从684.27 ℃上升至994.04 ℃，P3点热循环曲线的峰值温度从568.56 ℃上升至765.17 ℃.对P2点的计算所得热循环曲线和试验测得的热循环曲线较为吻合，结果表明多道复合焊的温度场计算结果较为合理，可以用于应力场计算.
3.3 应力场计算结果
试样S2的焊后残余应力分布如图7 ~ 图8所示，不同PL下焊件上表面位于1/2焊件长度处的残余应力分布规律如图9所示. 从图7 ~ 图8可以看出，纵向残余应力σx的高值拉应力出现在HAZ、焊缝中近1/2板厚处和打底焊缝，峰值拉应力(615.34 MPa)出现在打底焊缝；由于板厚较大、焊缝冷却速度较快，出现了较高的压应力，峰值压应力为−541.49 MPa. 横向残余应力σy在上、下表面的HAZ处均出现高值拉应力，峰值拉应力(607.76 MPa)出现在打底焊缝部位，高值压应力出现在焊缝两端的HAZ. 厚度方向应力σz在除了打底焊缝的其它地方都表现出较低的应力值，峰值拉应力出现在打底焊缝部位. 等效残余应力σVon的高值应力出现在HAZ和焊缝，峰值应力(573.61 MPa)出现在焊缝中近1/2板厚处.
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图 7 试样S2的残余应力分布Figure 7. Residual stress distribution in specimen S2. (a) longitudinal residual stress σx; (b) transverse residual stress σy; (c) pleate thickness direction σz; (d) equivalent residual stress σVon![]()
图 8 试样S2横截面的残余应力分布Figure 8. Residual stress distribution in the cross section of specimen S2. (a) longitudinal residual stress σx; (b) transverse residual stress σy; (c) plate thickness direction stress σz; (d) equivalent residual stress σVon![]()
图 9 残余应力在宽度方向的分布Figure 9. Residual stress distribution along the width direction. (a) longitudinal residual stress σx; (b) transverse residual stress σy; (c) platethickness direction σz; (d) equivalent residualstress σVon由图9可以看出，复合焊件上表面的σx和σVon应力水平较高，而σy和σz整体应力水平较低. 2.5 kW激光功率下，试验测得的σx和σy在数值上和计算所得的应力值有较好的一致性. σx在焊缝和HAZ表现为拉应力；σy在焊缝表现为压应力，在HAZ表现为拉应力；σz则在焊缝和HAZ表现出低应力的拉压交替. PL = 2.0 kW时，复合焊件中的整体应力水平最高. 当PL从2.0 kW增加到3.5 kW，上表面σx，σy，σz 的峰值拉应力和σVon均呈现下降趋势，σx ，σy ，σz 的峰值拉应力分别下降了110.1, 71.8和47.7 MPa. 当PL从3.5 kW增加到4.0 kW时，σx，σy ，σz和σVon的峰值应力均表现出一定的上升.
不同PL下复合焊件的残余应力峰值变化如表3所示. 激光功率PL在2.0 ~ 3.5 kW之间内，随着激光功率增加，σx，σy，σz和σVon各应力的峰值水平均表现为较小幅度的下降；当激光功率从3.5 kW增加到4.0 kW时，焊件中各应力分量的最大值有所上升.
表 3 残余应力峰值Table 3. Peak residual stresses激光功率
PL /kW纵向应力
σx/MPa横向应力
σy/MPa厚度方向应力
σz/MPa等效应力
σVon/MPa2.0 623.85 617.13 589.49 582.37 2.5 615.34 607.76 583.18 573.61 3.0 607.71 602.21 573.76 567.48 3.5 602.43 600.54 569.85 559.54 4.0 609.27 604.62 576.03 568.75 3.4 讨论
增加填充焊和盖面焊的激光功率PL，导致焊接热输入增大、焊件冷却速度降低，从而使焊缝和HAZ形成转变温度更高的显微组织和较粗大的晶粒. 因此盖面焊道中AF增加、GB和LB数量减少，电弧、激光各自作用区的CGHAZ中GB，AF数量增加，FGHAZ中QPF数量略有增加.
采用2.0 kW激光功率时，加热范围较小、焊件冷却速度较快，材料各部分温差较大，且已焊焊道重熔的金属较少导致焊件拘束度较大，故残余应力水平较高[11]；继续增加PL，材料的冷却速度下降、各部分温差下降，指状熔深增加、焊件拘束相对降低，因此应力有所降低；但当PL过大时，焊缝的塑性变形过大，反而会导致应力水平上升. 因此，随着PL增加，残余应力峰值表现出先降低再增加的规律. 因此，为获得较合适的显微组织类型和晶粒尺寸，同时尽量降低激光-MIG复合焊件的残余应力水平，可以选择激光功率3.5 kW、焊接速度0.5 m/min进行填充、盖面焊缝的焊接.
4. 结论
(1) 填充、盖面焊缝采用2.0 ~ 4.0 kW的激光功率时，X80管线钢激光-MIG复合焊接头中，CGHAZ的显微组织均为GB和LB，FGHAZ的显微组织主要为GB和QPF.
(2) 对不同激光功率下激光-MIG复合焊的焊接温度场和应力场进行了模拟计算，结果表明计算所得各焊道的熔池形貌、典型观测点的热循环曲线以及焊件表面的纵向、横向残余应力分布与实测结果较为接近.
(3) X80管线钢多道激光-MIG多道复合焊接的残余应力水平较高，激光功率为2.5 kW时，纵向残余应力、板厚方向应力在焊缝近1/2板厚处和打底焊缝中均出现高值的拉应力，横向残余应力峰值拉应力出现在打底焊缝处.
(4) 激光功率在2.0 ~ 4.0 kW之间时，随着激光功率的增加，等效残余应力峰值以及各应力分量的峰值表现出先下降再上升的规律. 当激光功率为3.5 kW时，复合焊接头的残余应力水平相对最低.
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图 1 不同溶液中Fe80B20和(Fe,Co,Ni,Cr)80B20的极化曲线
Figure 1. Polarization curves of the as-spun Fe80B20 and (Fe,Co,Ni,Cr)80B20 alloy ribbons in different solutions. (a) 3.5% NaCl; (b) 0.1 mol/L H2SO4; (c) 0.1 mol/L HCl
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图 2 Fe25Co25Ni25(B, Si)25金属玻璃磁滞回线
Figure 2. Hysteresis loops of Fe25Co25Ni25(B, Si)25 metallic glasses
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图 3 Cr，Mn，Fe，Co和Ni在不同基体中的扩散系数与温度的关系
Figure 3. Temperature dependence of the diffusion coefficients for Cr, Mn, Fe, Co and Ni in different matrices
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图 4 原粉的X射线图和DSC曲线
Figure 4. Corresponding X-ray patterns and DSC curves compared with the original powder. (a) X-ray diffraction pattern; (b) DSC curves；(c) local amplification of DSC curves
表 1 典型高熵非晶合金系
Table 1 Typical high-entropy amorphous alloy system
高熵非晶合金系 制备方法 临界尺寸d/mm Zr41.2Ti13.8Cu12.5Ni10Be22.5[24] 浇铸法 >50 Ti20Zr20Hf20Cu20Ni20[23] 铜模铸造法 1.5 Sr20Ca20Yb20Mg20Zn20[25] 铜模铸造法 4 Er20Tb20Dy20Ni20Al20[25] 铜模铸造法 2 Pd20Pt20Cu20Ni20P20[26] 熔渣包覆水淬法 10 Fe20Si20B20Al20Ni20[27] 球磨法 — Ti20Zr20Cu20Ni20Be20[28] 铜模铸造法 3 CoCrCuFeNiZr0.6[4] 熔体快淬法 — Fe25Co25Ni25(B，Si)25[29] 铜模铸造法 1.5 Er18Gd18Y20Al24Co20[30] 铜模铸造法 5 Fe25Co25Ni25Mo5P10B10[31] 熔体快淬法 1.2 (Fe1/3Co1/3Ni1/3)80(P1/2B1/2)20[32] 熔体快淬法 2 Fe46.8Mo22.7Cr13.6Co7.6C4.8B2.3Y1.2Si1.0[33] 激光熔覆 — 
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