Effect of beam oscillating and nitrogen alloying upon microstructure and mechanical properties in laser welding of molybdenum alloy
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摘要: 钼合金熔化焊接存在晶粒粗大、晶间偏析问题，导致接头力学性能差，采用激光束摆动和氮气合金化方法开展试验研究.结果表明，单独采用光束摆动措施后，焊缝区平均晶粒尺寸减小约28%，焊缝中心显微硬度从190 HV提高到200 HV，钼合金对接接头抗拉强度从29.83 MPa提高到130.03 MPa.单独采用氮气合金化(保护气体10 % N2 + 90 % Ar)的方法后，焊缝中心显微硬度从190 HV提高到240 HV，钼合金对接接头抗拉强度从29.83 MPa提高到350.94 MPa；在同时采用激光摆动与氮气合金化时接头抗拉强度达到了439.43 MPa，为母材抗拉强度的67.8%，且断裂模式由沿晶断裂转变为了沿晶断裂与穿晶解理断裂并存；分析认为氮气合金化对接头性能的强化效果得益于晶内和晶界处Mo2N相的生成.Abstract: The problems of coarse grains and intergranular segregation in molybdenum alloy welding lead to poor mechanical properties of the joints. The experimental study was carried out by using laser beam oscillation and nitrogen alloying. The results show that when using beam oscillating only, the average grain size in the weld zone was reduced by about 28 %, the microhardness of the weld center was increased from 190 HV to 200 HV, and the tensile strength of molybdenum alloy joint was increased from 29.83 MPa to 130.03 MPa. When using nitrogen alloying only (shielding gas 10 % N2 + 90 % Ar), the microhardness of weld center was increased from 190 HV to 240 HV, and the tensile strength of molybdenum alloy joint was increased from 29.83 MPa to 350.94 MPa. Furthermore, the tensile strength of the joint reached 439.43 MPa which was 67.8% of the tensile strength of the base metal when laser oscillating and nitrogen alloying were used simultaneously, and the fracture mode changes from intergranular fracture to intergranular fracture and transgranular cleavage fracture. The analysis shows that the strengthening effect of nitrogen alloying on the properties of the joint benefits from the formation of Mo2N phase in the grain and at the grain boundary.
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Keywords:
- molybdenum alloy /
- laser oscillation /
- grain refinement /
- nitrogen alloying
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0. 序言
拟建于新疆奇台的世界最大全向可转动射电望远镜(Qitai radio telescope, QTT)，其口径为110 m，为满足高频段观测要求，对天线精度提出了严苛的要求[1]. 轨道作为QTT天线的运动基准，其制造精度要求高[2-3]. 在实际轨道拼焊过程中，其单段轨道尺寸为4 000 mm × 600 mm × 200 mm，在实际生产中难于控制. 针对这类大型结构焊接变形问题，通常采用有限元计算方法进行焊接变形预测[4-6]. 然而采用有限元方法计算大型结构多层多道焊热-力耦合模型的计算量较大，QTT轨道拼接的单侧焊接包括15层80道焊缝，其模拟工作量较大. 针对这类问题，已有学者开展了厚板多层多道焊角变形的预测公式研究[7]. 但目前对角变形的研究一般仅考虑自由变形状态下的理论研究，而大型结构尺寸大，其重力对焊接接头弯矩作用明显，很可能影响焊接角变形量. 因此欲通过试验与有限元计算验证重力的影响，从而可以提高大型结构角变形预测精度.
1. QTT轨道缩比件焊接试验
试验件尺寸为单根轨道的八分之一，尺寸长为500 mm，宽75 mm，厚度30 mm. 坡口角度20º，坡口深度25 mm，间隙2 mm，对坡口进行6层总计10道多层多道焊试验. QTT采用42CrMo钢材，化学成分如表1所示.
表 1 42CrMo钢化学成分Table 1. Composition of 42CrMoC Cr Mo Si Mn S P Ni Cu 0.42 1.05 0.20 0.27 0.65 ≤ 0.035 ≤ 0.035 ≤ 0.030 ≤ 0.020 利用激光三角法测距法测量焊接变形，所用变形测量设备为PanasonicHG-C1200型微型激光位移传感器、8通道数据采集卡及基于LabView编写的数据采集界面.
2. QTT轨道缩比件有限元模型
缩比件模型节点总数为23 103个，单元总数为16 080个. 模型如图1所示，焊缝及附近采用过渡网格. 基于热-力耦合模型进行多层多道焊计算，热源选取双椭球热源模型，式(1)和式(2)分别为前半球和后半球热流密度[8].
$$ {q}_{{\rm{f}}}\left(x,y,{\textit{z}}\right)=\frac{6\sqrt{3}{f}_{{\rm{f}}}Q}{ab{c}_{{\rm{f}}}{\text{π}}\sqrt{{\text{π}}}}{{\rm{e}}}^{-3{x}^{2}/{a}^{2}}{{\rm{e}}}^{-3{y}^{2}/{ab}^{2}}{{\rm{e}}}^{-3{x{\textit{z}}}^{2}/{{c}_{{\rm{f}}}}^{2}} $$ (1) $$ {q}_{{\rm{r}}}\left(x,y,{\textit{z}}\right)=\frac{6\sqrt{3}{f}_{{\rm{r}}}Q}{ab{c}_{{\rm{r}}}{\text{π}} \sqrt{{\text{π}}}}{{\rm{e}}}^{-3{x}^{2}/{a}^{2}}{{\rm{e}}}^{-3{y}^{2}/{ab}^{2}}{{\rm{e}}}^{-3{x{\textit{z}}}^{2}/{{c}_{r}}^{2}} $$ (2) 式中：
$ {q}_{{\rm{f}}} $ 为前半球热流密度函数（J/mm3）；$ {q}_{{\rm{r}}} $ 为后半球热流密度函数（J/mm3）；a为双椭球半球宽度6 mm；b为双椭球加载深度3 mm；cf为前半球长度6 mm；cr为后半球长度12 mm. 42CrMo热物理性能参数如表2所示[9]. 装夹采用三点装夹方式，重力边界条件通过体积力加载，利用Marc非线性有限元分析软件进行分析.表 2 42CrMo材料热物理性能参数Table 2. Thermal-physical properties of 42CrMo温度T/℃ 比热C/(J·kg−1·℃−1) 热导率λ/(W·m−1·℃−1) 热膨胀系数α(10−6) 弹性模量E/GPa 屈服强度ReL/MPa 0 550 52.8 1 217 805 200 600 51.3 2.1 203 720 400 753 45.5 5 180 550 600 820 41.5 7 165 340 800 1 500 32.2 8.3 147 240 1 000 780 38.5 1.2 90 100 3. 缩比件结果对比
如图2外观照片可以观测到，10道焊后对接平板产生了明显的焊接角变形. 图3中对比了试验与数值模拟的角变形结果，模拟结果中包括考虑与不考虑重力边界的两种状态.
由图3可知，自由状态下有限元计算的角变形量约为70 mm，而试验测量值约为28.5 mm,两者结果差距较大. 而考虑重力边界的作用下，其角变形量明显减小，约为30.2 mm，与试验结果接近.
根据以上试验与有限元分析结果可知，考虑重力影响的有限元计算结果与试验结果基本吻合，更接近实际情况，因此重力边界在焊接角变形预测时不可忽略.
4. QTT天线轨道实际分段角变形预测
实际QTT天线轨道分段尺寸为弦长4 007.44 mm，轨道宽度600 mm，厚度200 mm. 其坡口尺寸如图4所示，坡口深70 mm，间隙3 mm，坡口角度20º. 将填料焊缝数为15层85道焊缝，如图4所示为坡口几何与网格模型.
文中将按照上述实际尺寸，依据已验证热力耦合有限元模型进行焊接角变形预测分析，同时结合有限元分析结果，进一步修正多层多道焊角变形预测公式.
4.1 自由状态角变形预测
文献[7]基于材料力学以及横向收缩力模型的角变形公式推导了自由状态多层多道焊角变形预测公式，即
$$ {\beta }_{i}=2{k}_{2}{\theta }_{A}=2{k}_{2}\frac{{M}_{0}l}{6EI}\;\;({i}\geqslant 2) $$ (3) $$ l = {k_1}\frac{{\Delta {B_i}}}{{{\varepsilon _s}}} $$ (4) 式中：
${M}_{0}={P}_{i}\cdot {{\textit{{\textit{z}}}}}_{i}$ 为焊缝收缩引起的角变形弯矩，$ {P}_{i} $ 为第$i$ 道焊缝在构件中引起的收缩力，$ {{\textit{z}}}_{i} $ 为当前焊道中心与已完成焊缝截面中心的距离；$ {\theta }_{A} $ 为挠曲变形转角;$ {\beta }_{i} $ 为焊接收缩角变形转角；l为焊缝塑性区宽度，与屈服极限有关；E为弹性模量；I为惯性矩；$ {\Delta B}_{i} $ 为横向收缩量. 文献[10]校核了$ {k}_{1} $ 刚度影响系数0.38；$ {k}_{2} $ 为焊缝形状系数，取0.48. 考虑焊缝尺度影响，式(3)则为式(5)和式(6).当前焊道未能填满该层时
$$ {\beta }_{i}=2{k}_{1}{k}_{2}\frac{{\delta }_{i}{\Delta B}_{i}}{{\delta }_{iy}}\left({\delta }_{y}+\frac{{\delta }_{i}}{2}\right) $$ (5) 当前焊道填满该层时
$$ {\beta }_{i}=2{k}_{1}{k}_{2}\frac{{\delta }_{i}{\Delta B}_{i}}{{\delta }_{iy}}\left({\delta }_{y}+{\delta }_{i}\right) $$ (6) 式中：
$ {\delta }_{i} $ 为当前焊道厚度；$ {\delta }_{iy} $ 为焊缝的计算厚度；$ {\delta }_{y} $ 为已完成焊缝厚度.通过角变形公式计算以及数值模拟计算得出前8道角变形量如图5所示. 结果表明，自由状态下，多层多道焊角变形预测公式与不添加重力边界有限元计算结果基本一致.
4.2 考虑重力角变形预测
式(1)角变形公式的推导过程中忽略重力抗弯矩的影响. 而对于大型结构QTT轨道结构，其分段尺寸较大，因此在进行公式推导时添加重力影响因素. 对角变形公式修正如下.
式(3)中，考虑重力影响时第一道焊缝总弯矩为
$$ {M}_{0}={P}_{i}\cdot {{\textit{z}}}_{i}-G\cdot \frac{L}{2} $$ (7) 式中：
$ G $ 为结构的重力，此时在焊缝累积弯矩克服重力弯矩时才开始发生明显焊接角变形，即满足条件式(8)后，后续焊缝角变形采用式(5)与式(6)计算.$$ \sum \limits_{i=1}^{m}{p}_{i}\cdot {{\textit{z}}}_{i}>G\frac{L}{2} $$ (8) 如图6所示重力弯矩影响下有限元模拟的焊接角变形，其中当焊缝道数较少时，焊缝累积收缩弯矩小于重力弯矩作用，其角变形量变化不明显，而随着焊缝道数增加，最终克服重力弯矩作用，角变形量逐渐增大. 图7所示为有限元角变形结果. 表3为修正公式与有限元分析的最终变形结果对比.
表 3 重力状态下角变形量对比Table 3. Comparison of deforming results under gravity类别 角变形$ \;\beta $/rad 厚度方向最大值 x/mm 有限元 0.727 61 修正公式 0.8 62 5. 结论
(1) 缩比件多层多道焊角变形测量试验有效的验证了热-力耦合有限元计算结果，考虑重力影响的分析结果与试验结果更吻合.
(2) 开展QTT天线实际分段自由状态下的多层多道焊有限元分析及角变形公式计算，有限元分析结果与公式预测结果相符.
(3) 在自由状态角变形公式基础上，增加重力因素，修正了预测公式；公式计算结果与考虑重力有限元分析结果吻合，证明了修正公式的有效性.
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图 2 激光振镜焊接试验装置图
Figure 2. Pictures of the experimental setup. (a) SPI fiber laser device; (b) galvanometer scanner; (c) Ar/N2 protection device
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图 3 摆动方式及拉伸试样(mm)
Figure 3. Oscillating mode and tensile specimen. (a) scanning path of the oscillating laser; (b) dimensions of the specimen used for tensile test
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图 4 焊接接头横截面组织形貌
Figure 4. Cross sectional microstructure of weld joints. (a) joint under Ar without laser oscillating; (b) joint under N2 without laser oscillating; (c) joint under Ar with laser oscillating; (d) joint under N2 with laser oscillating
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图 5 4种焊接接头显微硬度分布图
Figure 5. Microhardness distribution curves of joints under different conditions
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图 6 母材与4个焊接接头拉伸曲线
Figure 6. Tensile test curves of BM and joints under different conditions
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图 8 焊接接头横截面背散射电子衍射图(EBSD)
Figure 8. EBSD images of cross section of joints. (a) joint under Ar without laser oscillating; (b) joint under Ar with laser oscillating; (c) joint under N2 without laser oscillating; (d) joint under N2 with laser oscillating
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图 9 焊缝横截面晶粒尺寸统计图
Figure 9. Statistics on grain size in cross-section of FZ of joints. (a) joint under Ar without laser oscillating; (b) joint under N2 without laser oscillating; (c) joint under Ar with laser oscillating; (d) joint under N2 with laser oscillating
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图 10 拉伸断口的SEM观察结果
Figure 10. SEM bright images of tensile fractures. (a) joint under Ar without laser oscillating; (b) joint under N2 without laser oscillating; (c) joint under Ar with laser oscillating; (d) joint under N2 with laser oscillating
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图 11 焊缝区TEM观察结果
Figure 11. TEM bright image of fusion zone produced. (a) in 10 % N2 + 90 % Ar; (b) in 100 % Ar
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