Achieving High Strength-plasticity of Nanoscale Lamellar Grain Extracted from Gradient Lamellar Nickel

The main procedure of this work is clearly illustrated as the schematic image in Figure 1. First, the raw material was processed by UADR technique after heat treatment. Then, the UADRed macroscopic sample was characterized through some microstructural characterization techniques such as electron back scatter diffraction (EBSD), scanning electron microscope (SEM) and so on. Thus, the location of regions with different microstructure could be acquired, which provided the base of micropillar sample preparation in different regions. After different micropillars were fabricated by Focus ion beam (FIB) technique, in-situ compression tests would be performed to obtain mechanical behaviors of these micropillars. Finally, the TEM samples were processed from compressed micropillars through FIB technique, and then transferred to copper stage by using Pt deposition technique for further TEM characterization. More details and parameters are described in the following chapters.

The material used in this work is Ni with a purity of 99.9 wt.%, and its chemical composition (Table 1) is measured using energy dispersive spectrometer (EDS) analysis. A polycrystalline Ni rod with a diameter of 10 mm and a length of 100 mm was annealed at 1000 °C for 10 h in vacuum to encourage grain growth and homogenization, which was then subjected to UADR at ambient temperature.

The working principle of UADR is that working head exerts a certain amplitude of ultrasonic vibration on metallic material surface along normal direction and simultaneously feeds along the uniaxial direction, which induces normal stress to refine grains and shear stress to elongate grains on the metallic material surface [25‐27]. During the UADR processing, which is shown in Figure 2, the annealed rod sample rotates at a velocity of 250 r/min, while a steel ball penetrates into sample surface and rolls along axial direction at a velocity of 10 mm/min. The amplitude of ultrasonic vibration of steel ball ranges from 12 μm to 14 μm with vibration frequency of 17.5–18.8 kHz. The sample is considered as being treated by one pass as the tool tip rolls from one end of the rod surface to the other end. To increase the plastic strain, the sample was treated by five passes with loads of 80 N, 100 N, 120 N, 140 N and 160 N, respectively. Cooling oil was used for temperature control during processing.

After being polished, cross-sectional microstructure of the UADRed pure Ni rod (Figure 3(a)) was analyzed by EBSD technique at 20 kV acceleration voltage with a working distance of 18 mm. The SEM imaging of the samples was conducted using the Zeiss Crossbeam 340 FIB/SEM System at the voltage of 6 kV. Before in-situ compression tests, details of microstructure along the depth can be observed through EBSD images as shown in Figures 3(b)–(d). In addition, the TEM sheets were prepared from each compressed micropillar in the SEM/FIB chamber. After thinning process, the TEM sheet with the thickness around 100 nm was processed and characterized by the FEI Talos 200X TEM system with the accelerate voltage of 200 kV.

In-situ uniaxial micro-compression tests were performed in a Carl Zeiss CossBeam340 SEM using a Brukers-Hysitron PI88 PicoIndenter under displacement control. The cylinder indenter was equipped with a 5 μm diameter diamond flat-punch indenter, which was much larger than the 2 μm micropillars in this work to subtract the effect of indenter on deformation behavior. Micropillars were prepared by focused ion beam (FIB) technique at 30 kV with a final finishing current of 0.1 nA or 0.05 nA from each representative region of the UADRed Ni sample. The height-to-diameter aspect ratio of micropillars was kept at 2–2.3 to avoid buckling during compression. Diameters of micropillars were the average values of top and bottom diameters measured from the SEM images, in favor of alleviating the effect of approximately 3° taper of micropillars. Micropillars were compressed at a constant strain rate of 10⁻³ s⁻¹ to roughly 25% strain, followed by a 1 s holding before unloading. In order to ensure the reliability and repeatability of the tests, in-situ micro-compression tests were repeated 5 times for each region. The most representative curves of 5 times tests for each region were chosen to discuss the deformation behavior in the following chapters. The force and displacement data of compression test was acquired by a piezoelectric actuator in the capacitive transducer. SEM probe monitors the compression behavior of micropillars and automatically generates real-time videos at any time.

As mentioned above, considering the existence of a slight taper, the micropillar diameters was acquired by averaging the diameters of top and bottom of micropillars to calculate the engineering stress. Moreover, in order to avoid the impact of indenter and basement below the micropillar on elastic deformation data and acquire a relatively accurate measurement of strain, the equation of a cylindrical punch derived by Frick et al. [28, 29] was applied to calculate displacement \(x\),where \({x}_{meas}\) and \({F}_{meas}\) represent the measured displacement and force, respectively. \({D}_{t}\) and \({D}_{b}\) are the diameters of top and bottom of micropillar, respectively. \({\nu }_{i}\) and \({E}_{i}\) are the Poisson's ratio and Young's modulus of diamond flat-punch taken to be 0.07 and 1220 GPa, respectively, while \({\nu }_{b}\) and \({E}_{b}\) represent the Poisson's ratio and Young's modulus of Ni selected as 0.31 and 207 GPa, respectively. Once the displacement was corrected, the engineering stress-strain curves were obtained, and subsequently converted to true stress-strain curves by using the homogeneous deformation model (assuming no volume change during the deformation).

For micropillars with smooth stress-strain curves, since the determination of the exact value of yield stress is difficult due to a continuous transition from elastic to elastoplastic deformation regimes of the stress-strain response, flow stress obtained by a strain offset of 1% instead of a conventional 0.2% offset stress, is designated as yield stress [30], as illustrated in Figures 4‐6. For micropillars showing discontinuous flow behavior, the yield stress is identified as the stress at the first discrete displacement burst that occurred post-1% strain in each curve, and a threshold of \(\Delta \varepsilon\) ≥ 0.002 is used to identify the burst events [31], as exhibited in Figure 7(a). In this study, the strain hardening rate (SHR) is defined as the slope between the true stress–strain data at yielding point, \({\varepsilon }_{y}\), and 10% strain due to the strain bursts during the stress–strain response of the micropillars. If there is a strain burst at the nominal strain value, the peak stress in the burst is used.

The EBSD image of the UADRed Ni rod (Figure 3(a)) shows the gradient microstructure exhibiting unambiguous plastic deformation from top surface to 450 μm along the depth. The gradient lamellar Ni can be roughly divided into four regions according to their differences in microstructurally geometrical characteristics. The depth range of the three regions in plastic deformation zone ranging from top surface to 120 μm is respectively specified as 0–10 μm, 10–50 μm and 50–120 μm from top surface. However, the UADR technique [32, 33] would introduce a certain extent of plastic deformation on the surface of pure Ni rod, especially the largest plastic deformation on the outermost layer, which leads that the outermost layer is difficult to be resolved under the EBSD technique [34‐36]. Therefore, the microstructural details of plastic deformation regions from top surface to 6 μm are unable to be acquired through the EBSD image shown in Figure 3(a). As shown by EBSD observation in Figure 3(b), the representative microstructural characteristic in the region ranging from 6 to 10 μm contains near nanoscale lamellar (NNL) grains with average lamellar thickness of ~ 190 nm (shown in Figure 3(e)) and hence is referred to as the NNL region. The lamellar GBs are parallel to the rolling direction. The adjacent region of 10–50 μm from surface has severely deformed structure characterized by parallel lamellar grains named as the ultrafine lamellar (UFL) region. The elongated boundaries are also straight and flat lamellar GBs parallel to the rolling direction, which is shown in Figure 3(c). Thickness of the lamellar GBs were averagely measured as 360 nm (indicated in Figure 3(f)), and their average length is 1500 μm. The depth range of 50–120 μm presents submicron-sized grains and this range is referred to as the transition layer (TL) region, grains within TL region are elongated with an average boundary spacing of 645 nm (presented in Figure 3(g)) and length of 2 μm. As shown in Figure 3(d), the elongated boundaries of the UFGs are inclined to the rolling direction. Finally, as shown in Figure 3(a), the average grain size of equiaxed CGs is measured about 50 μm at the depth of 450 μm, and region deeper than 450 μm can be seen as the undeformed (UD) region.

To study different mechanical behaviors of each region, micropillars with diameter ~ 2 μm is fabricated to perform in-situ micro-compressive tests at the depth of 8 μm (NNL), 40 μm (UFL), 120 μm (TL), 450 μm (UD) from the treated surface, respectively.

From the true stress-strain curves of the NNL, UFL and TL micropillars, the NNL (Figure 4(a)) and the UFL micropillars (Figure 5(a)) were found to have a similar mechanical behavior, that is, they both deform elastically at the first stage and then get into a strain hardening stage. However, the yield stress (~ 705 MPa) and the SHR (~ 2826 MPa) of the NNL micropillar were both measured to be relatively high, while those of the UFL micropillar were measured to be 635 MPa (yield stress) and 2012 MPa (SHR), respectively. It is easy to understand that smaller lamellar grain thickness leads to higher yield stress in the NNL micropillar. Interestingly, the SHR of the NNL micropillar is still higher than the UFL micropillar, which is different from the trade-off in yield strength and the strain hardening capacity observed in other macro-scale researches [37‐40]. The excellent strain hardening capacity of the NNL micropillar may result from no slip bands penetrating through micropillar surface during compression thus maintaining homogeneous plastic deformation, which could be reflected from the smooth curve (Figure 4(a)) and the post-compression SEM image (Figure 4(c)) of the NNL micropillar. Compared with the true stress-strain curve of the NNL and the UFL micropillar, the stress-strain curves of the TL micropillar (Figure 6(a)) are relatively discontinuous and apparent strain-burst appeared during compression. Besides, multiple slip bands were obviously observed by comparing Figure 6(b) and (c). As measured from Figure 6(a), the TL micropillar has the yield stress of 390 MPa and the SHR of 1410 MPa, both are much lower than those of the NNL and UFL micropillars.

According to the EBSD image of cross-section microstructure, the average grain size of the UD region is 50 μm. Hence, the 2 μm diameter micropillar fabricated from the UD region is more likely to be a single grain. The true stress-strain curve for the UD region shown in Figure 7(a) exhibits the yield strength of 310 MPa, followed by abundant discrete ‘strain bursts’ as shown by drop in stress. Each burst is due to rapid displacement during loading [41‐43], which means the flat punch tip is out of contact with the micropillar top and meanwhile stress sharply decreasing. These bursts correspond to the visible slip bands on micropillar surface shown in Figure 7(c). These slip bands existing on compressed UD micropillar are quite deep and lead to shear fracture, with length almost surround the micropillar itself. It is found that the length of slip band corresponds to the amplitude (measured as the stress increment in ascending stage of each burst [44]) changes of strain bursts in true stress-strain curves of different micropillars. The shear fracture in compressed UD micropillar may be caused by the activation of many slip bands and then rapid penetration through the UD micropillar without the block from GBs and dislocations during compression [29, 45]. In other similar studies, the researchers have pointed out that the slip bands in the UD micropillar result from collective and avalanche-like motion of dislocations [46], which annihilate at the internal GBs or free surfaces of the micropillars [47‐50]. The SHR of the UD region was measured to be 433 MPa, which was obtained by linearly fitting the peak stresses at the nominal strain values from the true stress-strain curve shown in Figure 7(a). The weak strain hardening capacity of the UD region may arise from pure slip mode without intersection of several slip systems. The phenomenon resembles deformation behavior of the CG Cu micropillar studied by Hou et al. [51], and therefore low-density defects of the CGs in the UD region are presumed to be the underlying reason for low strain hardening capacity of the micropillar in the UD region.

As mentioned above, the NNL micropillar consists of grains with the smallest lamellar thickness compared to other micropillars fabricated from other regions, contributing to the most GBs to block the motion of slip band. Therefore, the highest yield strength occurs in the NNL micropillar, consistent with the classical Hall-Petch effect that the strength of metallic materials increases significantly with the decrease of grain size [56, 57].

During the in-situ compression tests, the whole deformation process was recorded and the SEM images of different micropillars were exhibited in Figure 9 at specific strains ~ 0%, 10 %, 20 % and after unloading, respectively. As shown in the SEM images in Figure 9(a)–(d), the NNL micropillar exhibits a relatively homogeneous deformation from top to bottom sections of the micropillar without appearance of any shear slip band during the whole compression test. Conversely, slip system has been activated at strain ~ 20%, where some kinks emerging on the surface of the UFL micropillar are observed during compressive deformation (Figure 9(g)). Moreover, the SEM images (Figure 9(j)) of the TL micropillar show multiple slip bands parallel to each other at strain ~ 10%, which are very short and shallow. Significantly, as shown in Figure 9(n)–(p), it is found that localized shear fractures are responsible for the failure mechanism in the UD region, while the micropillars in other regions simply exhibit deformation rather than shear fracture. From the SEM images of all micropillars after compression test, it is visible that appreciable lateral bulge took place in all samples, which implies that the micropillars experienced plastic deformation during compression. However, lateral bulge in the NNL micropillar is distributed more uniformly than in the other micropillar, indicating that the plastic strain is accommodated homogeneously by the whole NNL micropillar rather than localized on one or several slip systems in other micropillars. It is speculated that many different slip systems are required to be activated in each grain aiming to coordinate with deformation of their many adjacent grains in the NNL micropillar. Consequently, the activation of many different slip system needs extremely high applied stress, leading to the most obvious strain hardening phenomenon in the NNL micropillar. Also, these different slip bands block the slip activities from each other and distribute uniformly through the whole micropillar, resulting in no apparent strain localization to penetrate through the whole NNL micropillar or form shear fracture. Thus, the NNL micropillar exhibits not only high strength but also good plasticity compared to other micropillars.

For further exploring the deformation mechanisms for micropillars in different regions, the TEM characterization was adopted to observe dislocation changes of different post-compression micropillars. Figure 10(b) and (c) shows the TEM images of post-compression NNL micropillar, where the numerous lamellar GBs and extensive dislocation could be observed clearly. Therefore, the dislocation pile-ups in the NNL micropillar are improved due to existing lots of GBs and a great deal of dislocation interaction. Meanwhile, combined with no visible ‘strain-burst’ in strain-stress curves of the NNL micropillar, it is considered that no extensive dislocation annihilated on free surface of the micropillar. Additionally, the lamellar GBs are always perpendicular to the loading direction through the post-compression TEM images in the NNL micropillar. However, shear stress named critical resolved shear stress (CRSS) of a certain magnitude along the slip plane is required for the formation of slip, and the maximum resolved shear stress (MRSS) is oriented as 45° relative to the loading direction [48]. Therefore, the motion of many slip bands formed inside the NNL micropillar is severely blocked by perpendicular lamellar GBs, where much more stress is required to continue the generation and motion of slip bands, resulting in no visible slip band on sample surface during in-situ compression test. In summary, the 90° orientation with respect to the loading direction and the extensive dislocation pile-ups also contribute to the highest yield strength and the best strain hardening capacity of the NNL micropillar.

Comparing the TEM images of the UFL micropillar (Figure 10(e)) with the NNL micropillar (Figure 8(b)), it is evident that lamellar thickness is larger in the UFL micropillar with smaller yield stress. However, the lamellar GBs are still perpendicular to the loading direction. Figure 5(a)–(c) shows several slight slip bands corresponding to slight strain-burst in true stress-strain curve of the UFL micropillar, illustrating collective dislocations escaping on slip planes [58‐61]. The collective escape of dislocations probably because there is no enough lamellar GBs in the UFL micropillar to block the motion of slip bands and thus several slip bands penetrate through the whole micropillar, causing the annihilation of dislocations on sample surface. Therefore, lower strain hardening capacity is obtained from the UFL micropillar compared with the NNL micropillar.

It is noted from comparing Figure 10(a), (d) and (g) that the main differences between the TL micropillar and the above two micropillars are increasing lamellar thickness, and GBs are inclined to the loading direction in the TL micropillar. As the Hall-Petch effect mentioned above, the increasing lamellar thickness leads to weak strength of the TL micropillar. However, it is worth noting that the lamellar thickness range of the TL micropillar is large, and the largest value is ~ 1.95 μm, which means that mainly one lamellar GB existing in the sample with diameter of ~ 2 μm and height of ~ 4 μm. Therefore, the Hall-Petch effect could not be suitable for revealing the strengthening mechanism of the TL micropillar with only one lamellar GB. In this case, the weak strength of the TL micropillar may be due to lacking GBs to block the motion of dislocation and leading to abundant annihilation of dislocations on the sample surface. In addition, it is found that much fewer dislocations distribute in the TL micropillar compared to the NNL and the UFL micropillars, as shown in Figure 10(h). The much fewer GBs and dislocations leads to the rapid propagation of slip band in the TL micropillar, which means only one or few slip system is required to activate for maintaining plastic deformation. Thus, no interaction of many slip systems to block the further and rapid propagation of slip bands, resulting in slip bands penetrating through the whole micropillar and forming shear fracture rapidly. Moreover, the phenomenon that GBs are inclined to the loading direction indicates the lamellar GBs in the TL micropillar are affected by MRSS on 45° orientation. Specifically, the inclined lamellar GBs are conducive to form lots of slip bands and unable to effectively divert and prevent the propagation of slip bands in the TL micropillar, thus many slip bands propagate to the micropillar surface along the 45° orientation without any barrier. Consequently, abundant dislocations are annihilated on the TL micropillar surface. This is consistent with the existing abundant strain-burst observed from true strain-stress curve and visible slip bands parallel to each other on the surface of the TL micropillar. Therefore, these differences are mainly responsible for the weak strength and the poor strain hardening capacity of the TL micropillar.