**2. Material and Methods**

In this study, IN625 powder (EOS GmbH, Munich, Germany) with a chemical composition corresponding to UNS N06625 and ASTM B443 was used. An EOSINT M290 (EOS GmbH, Munich Germany) laser powder bed fusion system equipped with a 400 W ytterbium fiber laser and the EOS IN625\_Surface 1.0 Parameter Set (laser power ~300 W, scanning speed ~1000 mm/s, hatching space ~0.1 mm, and layer thickness ~40 μm) was employed to fabricate two types of specimens: 10 × 10 × 10 mm3 cubic specimens for microstructure evaluation, and 85 <sup>×</sup> 18 <sup>×</sup> 3 mm3 rectangular blanks for tensile testing (Figure 1a). The blanks were built in two directions relative to the build plate, as defined in Figure 1a, and have been referred to as vertical or horizontal (parallel or perpendicular to the build direction, respectively) throughout this paper. The chemical compositions of the IN625 powder, the as-built LPBF alloy, and the wrought annealed alloy (reference) are shown in Table 1.

**Figure 1.** (**a**) Rectangular blanks and cubic specimens on the build plate; (**b**) tensile specimens (dimensions in mm;\*-reference dimension) [24].

**Table 1.** Chemical composition of IN625 alloy (in wt. %).


Following the LPBF, the build plate with cubic and rectangular specimens was subjected to stress relief (SR) annealing at ~900 ◦C for 1 h (EOS recommendations), followed by forced air cooling (~1.5 ◦C/s) [10,33]. The SR treatments were carried out in a Nabertherm H41/N furnace under argon continuous flow (~15 L/min). Next, all the printed specimens were cut from the platform, using a reciprocated saw, and the rectangular blanks were machined by EDM (electrical discharge machining) to obtain the dumbbell-shaped tensile testing specimens shown in Figure 1b.

Finally, some SR specimens were reserved for future study, while the others were subjected to either hot isostatic pressing (HIP, Avure Technologies, Quintus QIH-3, Columbus, OH, USA) under pressurized argon atmosphere, followed by furnace cooling (~0.1 ◦C/s) [34–36]; or high temperature solution treatment (ST) for 1 h in an open-air furnace (Pyradia, Longueuil, QC, Canada), followed by air cooling (~0.5 ◦C/s) (Figure 2). Both the HIP and ST post-treatments were expected to homogenize the as-built LPBF microstructure and decrease the anisotropy of the IN625 alloy's mechanical properties. Since HIP is time- and resource-consuming, and can result in undesirable grain growth [16], ST is seen as its economic and technologically sound alternative.

**Figure 2.** Schematic presentation of the post-processing sequence of laser powder bed fusion (LPBF) alloy.

For this study, it was decided to carry out both treatments at the same temperature in the 1100–1150 ◦C range. This decision can be explained by the fact that HIP at this temperature had already been successfully used to reduce processing-induced porosity, and to homogenize and recrystallize LPBF-built microstructures [16,24,34]. On the other hand, for a high carbon version of IN625 alloy (~0.045 wt. %C), a typical ST temperature is also in the 1100–1150 ◦C range [37].

For reference, wrought IN625 alloy annealed at 980 ◦C (Table 1), provided by McMaster Corp. and corresponding to ASTM B443 (Grade 1, with a grain size of ~13 μm), was also characterized in the framework of this study. The Grade 1 (fine-grained) alloy was chosen because it possessed higher mechanical characteristics above 600 ◦C than the solution-treated Grade 2 (coarse-grained) alloy. The temperature- and time-dependent behaviors of the LPBF and wrought IN625 alloys were compared in this study from the perspective of their concurrent industrial use.

Tensile testing with a strain rate of 10−3·s−<sup>1</sup> was conducted at 25, 427, 538, 593, 649, 760, and 871 ◦<sup>C</sup> (68, 800, 1000, 1100, 1200, 1400, and 1600 ◦F) using an MTS 810 testing system equipped with an infrared radiant heating furnace. High temperature testing was realized under argon atmosphere at a flow rate of 5–18 L/h. Prior to tensile testing, specimens were heated at a heating rate of 1 ◦C/s and maintained at the test temperature for 10 min. The temperature was controlled using three K-type thermocouples in contact with the specimen surface and evenly distributed along its gauge length to control the uniformity of the temperature distribution. The strain was calculated using data provided by the LVDT (linear variable differential transducer) of the testing machine. After each treatment and for each testing temperature, the yield strength (YS corresponding to 0.2% offset strain), the ultimate tensile strength (UTS), and the elongation to failure (ε) were determined. For each experimental point, three specimens were tested, and the mean values of YS, UTS, and ε and their confidence ranges at a confidence probability of *p* = 0.95 were calculated.

Creep tensile testing was conducted at 760 ◦C (1400 ◦F) at 0.5, 0.7, and 0.9 of the YS with a loading rate of 10 N·s<sup>−</sup>1. The testing system, atmosphere, gas flow, and heating rates for the creep testing were identical to those of the elevated temperature tensile testing. Three tests were conducted for each creep condition, and the rupture time (τ), the fracture strain (ε), and the steady or secondary creep rates (. ε) were determined.

The fracture morphology and microstructure were analyzed using scanning electron microscopy (SEM, Hitachi TM3030 system and Hitachi SU8230 system equipped with an electron backscatter diffraction (EBSD) unit). The microstructural analysis was performed on the horizontal (XY) and

vertical (ZX) reference faces of the cubic specimens (Figure 1a). All the specimens were polished manually (down to 1 μm grit size), and then using a vibrometer and colloidal silica (0.05 μm grit size). For EBSD analysis, samples were tilted at 70◦ and scanned at 20 kV, with a step of 1–2 μm.
