Review on Regualtion Means in Wire Arc Additive Manufacturing of Aluminum Alloy
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摘要:
铝合金应用领域广，电弧增材制造(WAAM)成形大型结构件快，使铝合金电弧增材制造技术取得了长足的进步. 然而，使用 WAAM 工艺加工的铝合金零件易存在较多缺陷，成形质量也难以达到实际应用要求，常需要采用各种手段进行调控，以改善电弧增材制造铝合金的成形过程和组织性能. 针对铝合金电弧增材制造，本文总结了国内外过去十年来的研究情况，包括高水平论文发表数量和引用情况以及工程化应用等.针对电弧增材制造铝合金中存在的主要问题及其形成机理，从单一调控手段和复合调控手段两个角度进行综述，重点论述了各调控手段的优势和局限性，并对未来需要突破的研究方向进行了展望.
Abstract:Aluminum alloy possesses extensive applications, and wire arc additive manufacturing (WAAM) has emerged as a rapid method for producing large structural parts, leading to significant advancements in aluminum alloy wire arc additive manufacturing technology. However, the utilization of WAAM for manufacturing aluminum alloy parts often results in various defects, hampering the attainment of desired forming quality for practical applications. Consequently, diverse strategies are frequently employed to regulate and control both the forming process and microstructural properties of aluminum alloy produced via wire arc additive manufacturing. This study presents a comprehensive review of research conducted over the past decade on wire arc additive manufacturing of aluminum alloy, including process methodologies, advantages, disadvantages, the number of high-level scholarly publications, citation frequency, and engineering applications. The primary issues encountered in WAAM aluminum alloy production, along with their underlying formation mechanisms, are described and analyzed. Furthermore, domestic and foreign control methods addressing these challenges are reviewed from three perspectives: process optimization, heat treatment, and hybrid wire arc additive manufacturing technologies. The strengths and limitations of each regulation mean are emphasized, and prospective avenues for future research are outlined.
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Keywords:
- aluminum alloy /
- wire arc additive manufacturing /
- regulation mean
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0. 序言
激光是一种能量密度大、热输入精准可控的定向高能技术手段，将高能激光束作用在材料表面形成局部高温区，使材料熔化或蒸发，这种由激光辐照导致的材料烧蚀损失为激光切割等提供了新的工艺方法[1-4]，其在航空航天、汽车工业等领域受到了广泛关注[5-6].在激光辐照范围内，母材随时间、空间会发生急剧变化，涉及多物理场耦合的复杂传热传质过程.气流环境、工件所处空间位置等外界因素的存在在影响材料去除强度、改变烧蚀速率的同时，进一步加剧了激光烧蚀过程的复杂性和瞬态性[7-11].目前，激光、材料以及周围环境之间的复杂相互作用仍是提高激光切割等工艺效率研究过程中亟需探索的科研难题.
国内外学者针对气流、空间位置等外界环境对激光烧蚀过程的影响进行了广泛地研究.Robin等人[12]研究了超音速横向气流对高功率激光照射不透明固体表面的影响，气流在熔体表面施加剪切应力，导致熔体在区域后缘去除，增加了熔体/固体边界的衰退率. 当静止靶材表面与水平方向成一定夹角时，重力效应可使熔透时间显著缩短[13].王伟平等人[14]发现切向气流和靶面之间的换热量与激光参数、气流参数以及靶参数均相关，气流与靶面的对流换热使激光加热效应所需的能量增加.Qin等人[15]提出了一种激光束传播方向与重力方向相反的向上打孔法，将其与向下打孔法进行对比，发现在重力作用下，材料去除质量提高了2倍以上，且损伤范围增大，提高了打孔效率.袁春等人[16]对比研究了不同气流环境下的激光辐照效应，研究发现，在切向空气气流环境下，氧化反应可促进钢靶温度升高，烧蚀效果最显著；在氮气气流环境下，钢靶有一定的烧蚀；在自然对流环境下，烧蚀效果最差.Wang等人[17]观察到气流的机械侵蚀提高了烧蚀速率，并使烧蚀面变得光滑，其对流冷却效应延缓了背表面温度响应，改变了激光光斑边缘的烧蚀形貌和成分.Yuan等人[18]分别进行了单片高温合金板、未填充夹层板、隔热材料填充夹层板和烧蚀材料填充夹层结构的激光烧蚀试验.结果表明，在高超声速横向气流作用下，烧蚀材料填充夹层结构的抗激光烧蚀性能明显优于其他结构.
合理的数值模型对于研究激光−材料−周围环境的相互作用过程，分析激光烧蚀过程中的温度分布、烧蚀形貌等具有重要意义.杜太焦等人[19]模拟研究目标在激光辐照下的高速运动过程，发现切向气流环境下的对流换热在一定程度上减弱了激光辐照的温升效应.Chen等人[20]对圆柱形密封爆炸装置的激光辐照响应进行了数值模拟，结果表明，与无气流相比，切向气流可显著缩短起爆时间.Huang等人[21]研究发现超音速气流与烧蚀坑不同位置之间的不均匀对流换热作用改变了烧蚀坑的对称形态.随着烧蚀坑的加深，流动模式由闭坑流向开坑流转变，进而显著改变局部流动特性，影响激光烧蚀行为.Wang等人[4]发现在激光切割过程中，熔化材料在重力和辅助气流的影响下向下移动，并将热量传递给底层材料以进一步熔化.Xing等人[22]提出了一个热流固耦合模型，研究亚音速切向气流作用下激光辐照金属板的耦合变形和动态破坏过程.研究发现，气流不仅加速了变形，而且使塑性应变和破裂模式变得不对称.Ma等人[23]建立了一个热−流体−烧蚀耦合数值模型，模拟高速风洞环境下的激光烧蚀，结果表明，激光烧蚀深度与激光功率密度、气流速度等有关.
目前，激光烧蚀过程中的质量损失主要是基于辅助气流展开研究，如不同性质气流环境下的温升差异、不同强度气流环境下的烧蚀速率、不同材料的抗烧蚀性能等方面，很少涉及工件所处空间位置对熔体去除过程的影响，更加缺乏气流与空间位置联合影响诱导熔体去除的研究.文中利用有限元分析方法，明确考虑了气流和所处空间位置的联合作用，建立了能够反映激光能量、表面气流、空间位置和金属材料之间相互作用的数值耦合模型，直接模拟了不同空间位置条件下，亚音速切向气流与大尺度熔池的相互作用过程，结合高速摄像技术，对比研究了空间位置对激光烧蚀过程中气流流动行为及熔体剥离损失的影响，以期对相关领域的研究提供试验数据和理论支撑.
1. 数值模型建立
文中采用有限元软件COMSOL Multiphysics对不同空间位置的亚音速切向气流与大尺度熔池之间的相互作用过程进行了模拟计算，解释了气流和工件所处空间位置等环境因素对激光辐照效应的影响.研究选用低碳钢作为母材，计算过程中所采用材料的主要热物理性能参数见表1.
密度
ρ/(kg·m−3)动态粘度
$ \mu $/(kg·m−1·s−1)固相线
Ts/K液相线
Tl/K热导率
k/(W·m−1·K−1)比热容
Cp/(J·kg−1·K−1)对流换热系数
h/(W·m−2·K−1)熔化潜热
H/(105 J·kg−1)7 860 0.005 5 1 722 1 754 61.1 745 40 2.7 1.1 激光辐照
在高能激光的辐照下，金属表面温度升高而熔化，假设激光入射方向始终与金属表面相垂直，激光能量呈高斯分布，其热流密度随空间位置分布的函数为[28-30]
$$ I = \frac{{2P}}{{\text{π} {R^2}}}\exp \left( {\frac{{ - 2{r^2}}}{{{R^2}}}} \right) $$ (1) 式中：P为激光功率；R为激光光斑半径；r为金属表面某一点到激光光斑中心的距离.
1.2 热量传输
在激光辐射范围内，母材温度随时间、空间发生急剧变化，可根据能量方程对热量传输过程进行求解[31]，能量方程为
$$ \rho {C_{\mathrm{p}}}\left( {\frac{{\partial T}}{{\partial t}} + v \cdot\nabla T} \right) = \nabla \cdot \left( {k\nabla T} \right) + I $$ (2) 式中：$ \rho $，$ {C_{\mathrm{p}}} $，$ k $分别为材料的密度、比热容和热导率；T为温度；t为时间；在固体域内，$ v $为模型部件运动的速度，在该模型中固体域或热源均无移动，因此$ v $ = 0；在流体域中，$ v $为流体速度.
模型中，假设熔体最高温度为
3034 K[32]，即材料的蒸发温度，仅考虑对流换热造成的能量损失，不考虑材料的蒸发损失及高速气流带来的气动热，则作用于自由表面的能量通量为$$ k\nabla T = I - h\left( {T - {T_0}} \right) $$ (3) 式中：h为对流换热系数[21]；T0为环境温度，设置为293.15 K.
对于其他边界
$$ k\nabla T = - h\left( {T - {T_0}} \right) $$ (4) 1.3 流体流动
将空气及熔池中金属液体视为不可压缩的牛顿流体，其流动由连续性方程和Navier-Stokes方程控制的质量和动量运输描述[33-34]，连续性方程和Navier-Stokes方程分别为
$$ \nabla \cdot v = \frac{m}{\rho } $$ (5) $$ \rho \left( {\frac{{\partial v}}{{\partial t}} + \left( {v \cdot \nabla } \right)v} \right)=- \nabla p + \nabla \cdot \left( {\mu \left( {\nabla v + \nabla {v^T}} \right)} \right) + \rho g + $$ $$ \sigma \kappa n + {S_{\mathrm{M}}} $$ (6) 式中：m为质量源项；p为压力；$ \mu $为动态粘度；g为重力加速度；$ \sigma $为表面张力系数；$ \kappa $为气液界面曲率；$ n $为界面的单位法向量；SM为动量源项，即导致熔体剥离的气流的机械侵蚀力.
采用相场法追踪空气、液态金属两不相混合流体的界面[35-36]，COMSOL求解公式为
$$ \frac{{\partial \phi }}{{\partial t}} + v \cdot \nabla \phi = \nabla \cdot \left( {\frac{{\gamma \lambda }}{{{\varepsilon ^2}}}} \right) \cdot \nabla \psi $$ (7) $$ \psi = - \nabla \cdot {\varepsilon ^2} \cdot \nabla \phi + \left( {{\phi ^2} - 1} \right) \cdot \phi $$ (8) 式中：$ \phi $为相场变量，模型的空气域中设置为$ \phi = - 1 $，液态金属域中设置为$ \phi = 1 $，分别对应于体积分数$ {V_{\mathrm{f}}} $ = 0和1，其中$ {V_{\mathrm{f}}} = \left( {1 + \phi } \right)/2 $.两相之间的接触角设置为$ \text{π} $；$ \lambda $为混合能量密度；$ \varepsilon $为界面厚度控制参数；$ \gamma $为界面迁移率.
表面张力系数与混合能量密度、界面厚度的关系为
$$ \sigma = \frac{{2\sqrt 2 \lambda }}{{3\varepsilon }} $$ (9) 界面迁移率$ \gamma $决定了Cahn-Hilliard扩散的时间尺度，必须足够大以保持恒定的界面厚度，但仍应保持足够小以使对流项不会被过度抑制，其与界面厚度控制参数$ \varepsilon $与迁移率调整参数$ \chi $密切相关，关系式为
$$ \gamma = \varepsilon {\chi ^2} $$ (10) 1.4 网格划分
网格的划分对保证模型的收敛性和准确性至关重要.激光烧蚀过程中，激光辐照及气固界面附近区域的温度、速度等数据变化梯度大，网格划分较为细密，最小网格单元尺寸设置为0.036 mm，其他区域网格划分较为粗糙，如图1所示，计算域尺寸设置为55 mm × 18 mm，其中材料厚度为6 mm.在网格随固液界面移动而变形过程中，为了防止其过度扭曲，设置一个最小单元质量的值作为网格质量的判据，当单元质量低于此值时，自动对网格进行重新划分.
在图1中，气流沿金属表面切向流入，设为入口边界，空气域其他边界设置为出口边界，出口压力设置为101.325 kPa. 图1中$ \alpha $为金属钢板受激光辐照的表面与水平表面的夹角，钢板的空间位置随$ \alpha $的变化而变化，其值越大，重力沿金属受激光辐照表面的分量越大.调整金属钢板的空间位置使得$ \alpha $分别为90°，60°，30°，采用全耦合方法对不同空间位置的表面气流作用下的激光烧蚀过程进行数值求解.
2. 模型验证与分析
图2为气流作用下激光烧蚀装置系统示意图，系统由光纤激光器、机器人、气流喷嘴、高速摄像机、激光照明器等设备组成，可以实现金属钢板在不同空间位置下的激光烧蚀.在试验过程中，激光入射方向与钢板表面始终保持垂直，以约为40 m/s的气流速度通过位于熔池前方的气流喷嘴，并沿钢板表面切向流入，通过调整钢板空间位置、气流作用方向及激光入射方向，依次开展了$ \alpha $为90°，60°，30°的切向气流作用下激光烧蚀试验.利用高速摄像机，设定
5000 帧/s的采样频率捕捉熔池表面波动和熔体剥离飞溅图像，同时，利用激光照明器为熔池图片采集提供足够的亮度，从而对激光烧蚀全过程进行实时监测.基于钢板在不同空间位置下的激光烧蚀试验观测结果，量化金属烧蚀损失所形成熔坑的尺寸，测得熔坑长度a、熔坑宽度b、熔坑深度c，将熔坑近似为半个椭球体，进而可计算其体积为
$$ V = 1/2 \times 4\text{π} \left( {a/2} \right)\left( {b/2} \right)c/3 $$ (11) 在不同空间位置(试验1：$ \alpha $ = 90°；试验2：$ \alpha $ = 60°；试验3：$ \alpha $ = 30°)条件下，金属钢板激光烧蚀熔坑体积的变化趋势如图3所示. 调整钢板空间位置至$ \alpha $ = 90°，熔体剥离飞溅可形成体积为271.1 mm3的熔坑；当$ \alpha $减小到60°，熔坑体积从271.1 mm3减小到176.1 mm3，比$ \alpha $ = 90°时减小35.04%；当$ \alpha $进一步减小到30°时，熔坑体积减小到130.4 mm3，比$ \alpha $为60°时减小25.95%. 分析发现，当调整钢板空间位置使$ \alpha $逐渐减小时，熔坑体积随之减小，即烧蚀过程中金属质量损失降低，且金属质量损失下降幅度随$ \alpha $的进一步减小有所减缓.
在$ \alpha $为90°时，使用高速摄像机拍摄获得激光烧蚀过程中熔池表面形态演变及熔体剥离飞溅的图像，如图4所示，将其与如图5所示的数值模拟结果进行比较.对比发现，试验与模拟所获得熔体流动及剥离过程基本一致. 在$ \alpha $为90°时，重力作用方向平行于金属受激光辐照表面，在气流侵蚀和重力作用下，熔池表面熔体发生明显波动形成凸起，如图4(a)和图5(a)所示.随后，越来越多的熔体向下移动，在熔池外形成液柱，如图4(b)和图5(b)所示.在气流和重力的持续作用下液柱被拉伸，当液柱变得足够细时，熔体克服表面张力剥离熔池形成飞溅，如图4(c)，4(d)和图5(c)，5 (d)所示.此外，对比图4(d)所示的试验结果及图5(d)所示的模拟结果，可以发现试验所获得的熔坑长度为9.94 mm，模拟所获得的熔坑长度为10.04 mm，误差仅为1%，二者吻合良好. 以上分析表明所建立的数值模型能够较为准确地反映气流和工件所处空间位置对熔体剥离损失的影响.
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图 4 高速摄像机拍摄获得$ \alpha $ = 90°时激光烧蚀过程Figure 4. Laser ablation process at $ \alpha $ = 90° captured by a high-speed camera. (a) t0 + 0 ms; (b) t0 + 10.9 ms; (c) t0 + 13.3 ms; (d) t0 + 14.7 ms![]()
图 5 数值模拟获得$ \alpha $ = 90°时激光烧蚀过程Figure 5. Laser ablation process at $ \alpha $ = 90° obtained by numerical simulation. (a) t0 + 0 ms; (b) t0 + 10.9 ms; (c) t0 + 13.3 ms; (d) t0 + 14.7 ms3. 结果与讨论
图6至图11分别展示了$ \alpha $为90°，60°，30°时0.3 Ma切向气流作用下，激光烧蚀过程中气流及熔池流动行为的模拟结果.
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图 6 $ \alpha $ = 90°激光烧蚀过程中气流流动行为及熔体剥离过程的模拟结果Figure 6. Simulation results on flow behavior of airflow and melt peel-off process during laser ablation at $ \alpha $ = 90°在图6、图7的5.76 ms时，气流沿金属右侧表面从上方流入，到达熔池上边缘时，以2 ~ 11 m/s的速度作用在熔体表面并推动其向下运动.在图6、图7的9.47 ms时，在气流和重力的联合作用下，熔体被加速且速度分布较为均匀，大量熔体同时获得向下流动的高动量，在金属表面形成约2.4 mm高的凸起.随后，凸起继续向下移动，在表面张力作用下，熔体并未剥离熔池，而是在熔池外形成一个与金属表面呈较大夹角$ \theta $的倾斜液柱，如图6、图7的10.56 ms所示.气流及重力作用方向与液柱几近垂直，在气流与重力的持续作用下，液柱逐渐被拉伸形成仅约0.45 mm厚的细长液柱，且液柱中熔体被加速至1.6 m/s以上，随后高速熔体克服表面张力剥离熔池形成飞溅，如图6、图7的12.27 ms，13.82 ms所示.上述熔体剥离过程表明，须同时满足熔体速度足够大和液柱足够细两个条件才能使得熔体剥离熔池.
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图 7 $ \alpha $ = 90°激光烧蚀过程中熔体流动行为模拟结果Figure 7. Simulation results on flow behavior of melt during laser ablation at $ \alpha $ = 90°将图6、图7与图8、图9进行对比，可以看出，当$ \alpha $减小为60°时，重力沿金属受激光辐照表面的分量减少，熔体的加速进一步依赖于气流的侵蚀作用，因而较$ \alpha $为90°时熔体流速分布较为不均匀，越靠近凸起尖端，熔体流速越大，熔池底部熔体向下流动速度减慢，波动熔体体积减小，金属表面形成约2.1 mm高的凸起，比$ \alpha $为90°时降低12.5%，如图8、图9的5.90 ms所示.这也使得波动熔体对气流的阻碍作用减弱，气流偏转角度减小，进而导致凸起被推出熔池后形成的倾斜液柱与金属表面的夹角$ \theta $减小，如图8、图9的7.78 ms所示.在图8、图9的10.10 ~ 11.00 ms中，液柱熔体被加速至1.5 m/s以上，比$ \alpha $为90°时降低了6.25%，当液柱被拉伸至约0.35 mm厚时，熔体实现剥离.虽然液柱熔体速度较$ \alpha $为90°时低，但其较小的体积使得剥离更早发生.同时由于气流偏转角度较小，液柱被进一步拉伸时更加靠近金属表面.
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图 8 $ \alpha $ = 60°激光烧蚀过程气流流动行为及熔体剥离过程的模拟结果Figure 8. Simulation results on flow behavior of airflow and melt peel-off process during laser ablation at $ \alpha $ = 60°![]()
图 9 $ \alpha $ = 60°激光烧蚀过程熔体流动行为模拟结果Figure 9. Simulation results on flow behavior of melt during laser ablation at $ \alpha $ = 60°将图8、图9与图10、图11进行对比，当$ \alpha $减小为30°时，重力沿金属受激光辐照表面的分量进一步减少，熔体的加速几乎全部依赖于气流的作用，波动熔体在金属表面形成仅约1.4 mm高的凸起，比$ \alpha $为60°时降低33.3%，如图10、图11的5.70 ms所示.更小体积的波动熔体被推出熔池，且形成的倾斜液柱与金属表面的夹角$ \theta $进一步减小，如图10、图11的7.79 ms所示.在图10、图11的10.01 ms中，液柱熔体被加速至1.4 m/s以上，比$ \alpha $为60°时降低了6.67%，这使得熔体需要为剥离熔池积累较长时间的动量，因此相比于$ \alpha $为60°时剥离时刻没有明显提前，如图10、图11的10.83 ms所示.
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图 10 $ \alpha $ = 30°激光烧蚀过程中气流流动行为及熔体剥离过程的模拟结果Figure 10. Simulation results on flow behavior of airflow and melt peel-off process during laser ablation at $ \alpha $ = 30°![]()
图 11 $ \alpha $ = 30°激光烧蚀过程中熔体流动行为模拟结果Figure 11. Simulation results on flow behavior of melt during laser ablation at $ \alpha $ = 30°图12和图13总结了表面气流和空间位置联合作用下熔体的剥离过程.在图12中，在气流和重力的联合作用下，大量熔体获得向下流动的高动量，移出熔池后形成一个与金属表面呈大夹角$ \theta $的倾斜液柱，这使得气流及重力在很大程度上作用于液柱表面，液柱被沿金属表面拉伸，当液柱被拉伸得足够细长时，获得高动量的液柱熔体克服表面张力剥离熔池形成飞溅，如图13所示.随着$ \alpha $的减小，重力沿金属受激光辐照表面的分量减少，熔体的加速逐渐依赖于气流的侵蚀作用，液柱熔体流速减慢，且所形成液柱与金属表面的夹角$ \theta $的减小使得气流和重力对液柱表面的作用分量减少，从而导致熔坑尺寸减小、金属损失降低.
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图 12 熔体剥离前气流和空间位置联合作用示意图Figure 12. Schematic of the combined action of airflow and spatial position before melt peeling off![]()
图 13 熔体剥离后气流和空间位置联合作用示意图Figure 13. Schematic of the combined action of airflow and spatial position after melt peeling off4. 结论
(1)建立了一个考虑亚音速切向气流和工件所处空间位置联合作用的激光烧蚀数值模型
(2) 调整金属空间位置使$ \alpha $从90°减小到30°，重力沿金属受激光辐照表面的分量进一步减少，熔体的加速逐渐依赖于气流的侵蚀作用，液柱熔体流速由1.6 m/s减慢至1.4 m/s.当$ \alpha $减小到60°时，较小体积的液柱使得剥离更早发生.当$ \alpha $进一步减小到30°时，低速熔体需积蓄较长时间的动量，剥离时刻提前不明显.
(3) 工件所处空间位置直接影响烧蚀损失.调整金属钢板空间位置，随着$ \alpha $的减小，气流和重力对液柱表面的作用分量减少，激光烧蚀所形成的熔坑尺寸减小，熔池剥离程度和金属质量损失降低.
(4) 气流和金属所处空间位置联合作用诱导熔体去除主要表现在两方面：一方面，熔体获得向下流动的高动量，移出熔池后形成倾斜液柱，液柱熔体被加速；另一方面，液柱被拉伸变得细长，使得熔体易于克服表面张力，最终剥离熔池形成飞溅.
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图 1 铝合金WAAM技术研究团队分布图
Figure 1. Research team distribution map of aluminum alloy WAAM technology
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图 2 铝合金WAAM研究领域的国家占比
Figure 2. The proportion of national proportion in the field of WAAM aluminum alloy
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图 3 铝合金WAAM主题的已发表文章数量和被引频次
Figure 3. The number of published articles and the frequency of citations on aluminum alloy WAAM topics
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图 4 铝合金电弧增材制造技术在航空航天工业中的应用
Figure 4. Examples of wire arc additive manufacturing technology of aluminum alloy in the aerospace industry: (a)Fuel tank structure (b)Rocket engine nozzle (c)Rocket propellant pressure tank (d)The Dragon III launch vehicle
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图 5 英尼格玛使用电弧增材制造打印的铝合金汽车底盘
Figure 5. aluminum alloy car chassis made by Enigma using wire arc additive manufacturing
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图 8 不同保护气体成分下铝合金构件的外观形貌和性能 [16]
Figure 8. Morphology and mechanical properties of aluminum alloy components with various shielding gas composition
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图 9 不同保护气体流速下的孔隙率[17]
Figure 9. Porosity of aluminium alloys with various gas flow rates
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图 10 不同层间温度下
5356 铝合金的孔隙率[18]Figure 10. Porosity of
5356 aluminum alloy under different interlayer temperatures![]()
图 11 不同焊接速度下Al-Mg铝合金的力学性能[21]
Figure 11. The mechanical properties of Al-Mg aluminum alloy with different travel speed
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图 12 原材料和制造样品[23]
Figure 12. Raw materials and fabricated samples:
(a) TiC nanoparticle-enhanced weld wire; (b) conventional butt weld; (c) single layer overlay; (d) 3D printed multi-layer
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图 14 试样在不同状态下的电子背散射衍射图和晶粒尺寸分布统计图[27]
Figure 14. Electron backscattered diffraction (EBSD) graphs and statistical diagrams of grain size distribution of samples in different states：(a-b) the as-deposited samples; (b-d) the T6 heat-treated samples
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图 15 沉积态试样和热处理试样的3D气孔图[28]
Figure 15. 3D views of micropores of as-deposited and heat-treated samples
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图 16 WAAM + 层间FSP增材制造技术示意图[29]
Figure 16. Schematic diagram of WAAM + interlayer FSP additive manufacturing
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图 17 采用WAAM + 层间FSP方法制备的Al-Si合金的孔隙分布 [30]
Figure 17. Pores distribution in the WAAM + interlayer FSP Al-Si alloy
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图 18 LSP系统示意图及LSPed样品近表面区域塌陷孔隙的SEM图[33]
Figure 18. Schematic diagram of an LSP system and SEM image of collapsed pores in the near-surface zone of LSPed sample
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图 19 基于超声冲击强化的电弧增材复合制造[34]
Figure 19. Hybrid arc additive manufacturing based on UIT technology
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图 22 层间锤击设备示意图及实物图[43]
Figure 22. Schematic diagram and physical view of interlayer hammering equipment
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图 24 基于外部交变磁场的电弧增材复合制造[46]
Figure 24. The schematic diagram of WAAM process with external alternating magnetic field
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图 25 基于微合金化 + 热处理的电弧增材复合制造[49]
Figure 25. Hybrid arc additive manufacturing based on micro-alloying and heat treatment
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图 26 基于激光冲击强化 + 热处理的电弧增材复合制造[51]
Figure 26. Hybrid arc additive manufacturing based on laser shock peening and heat treatment
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图 28 基于路径优化 + 激光冲击强化的电弧增材复合制造[53]
Figure 28. Hybrid arc additive manufacturing based on path optimization and laser shock peening
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