Applied Science, Faculty of
Engineering, School of (Okanagan)
DSpace
UBCO
Paul, Christopher G. M.
Master of Applied Science - MASc
University of British Columbia
Additive manufacturing (AM) technologies are invaluable for the manufacturing industry by enabling the repair of high-value components. Utilizing AM technology such as laser-directed energy deposition (L-DED) to repair casting inserts can increase tool life, reduce metal waste, and lower operating costs. Typically, L-DED systems employ a 1–3 mm diameter laser spot to melt feedstock material. However, recently developed micro-spot L-DED systems, if optimized correctly, could enhance repair applications by creating smaller beads and minimizing heat input. This thesis aims to define suitable parameters for a micro-spot L-DED system and apply these parameters to a proof-of-concept repair process for a high-pressure die casting (HPDC) insert. By varying main L-DED printing parameters, the deposition behaviour and process parameters for achieving high-density maraging steel were determined using a micro-spot L-DED system. Analyzing the aspect ratio and dilution relationship of beads defined a processing window for the ideal bead shape. Volumetric samples were then printed, and process parameters which created high-density material (>99%) were selected. Compared to larger laser diameters, the 0.3 mm laser significantly influences bead formation characteristics, weld pool stability, part shape accuracy, and internal defects. Internal defects such as pores and titanium oxide inclusions reduce tensile strength and ductility for this material. Variations in grain structures found throughout the deposited material were correlated with changes in cooling rate during L-DED processing. The developed L-DED repair process for the damaged HPDC insert removes the heat checking crack and deposits material with micro-spot L-DED, restoring the component’s geometry. After the repair process, external and internal analyses provide valuable insights into the geometrical accuracy of the repair region, the microstructural and microhardness differences iii between the softer deposited material, the harder base material, and the traits of the heat-affected zones. This work showed that micro-spot L-DED is feasible for repair, but it also showed key areas in the repair methodology that require improvement before industrial implementation. Additionally, the unique response of weld beads and cuboid samples to process parameters warrants further investigation, focusing on the optimization of gas flow and powder feed rate settings for micro-spot L-DED systems.
Performance and Repair Applications of 18Ni-300 Maraging Steel Printed with Micro-spot Laser Directed Energy Deposition by Christopher G. M. Paul B.A.Sc., The University of British Columbia, 2022 A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIRMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE in THE COLLEGE OF GRADUATE STUDIES (Mechanical Engineering) THE UNIVERSITY OF BRITISH COLUMBIA (Okanagan) August 2024 © Christopher G. M. Paul, 2024 i The following individuals certify that they have read, and recommend to the College of Graduate Studies for acceptance, a thesis entitled: Performance and Repair Applications of 18Ni-300 Maraging Steel Printed with Micro-spot Laser Directed Energy Deposition submitted by Christopher G. M. Paul in partial fulfillment of the requirements of the degree of Master of Applied Science. Dr. Michael Benoit, University of Waterloo, Mechanical and Mechatronics Engineering Supervisor Dr. Lisa Tobber, School of Engineering Co-Supervisor Dr. Alexander Uhl, School of Engineering Supervisory Committee Member Supervisory Committee Member University Examiner Dr. Ian Foulds, School of Engineering Dr. Christopher Collier, School of Engineering ii Abstract Additive manufacturing (AM) technologies are invaluable for the manufacturing industry by enabling the repair of high-value components. Utilizing AM technology such as laser-directed energy deposition (L-DED) to repair casting inserts can increase tool life, reduce metal waste, and lower operating costs. Typically, L-DED systems employ a 1–3 mm diameter laser spot to melt feedstock material. However, recently developed micro-spot L-DED systems, if optimized correctly, could enhance repair applications by creating smaller beads and minimizing heat input. This thesis aims to define suitable parameters for a micro-spot L-DED system and apply these parameters to a proof-of-concept repair process for a high-pressure die casting (HPDC) insert. By varying main L-DED printing parameters, the deposition behaviour and process parameters for achieving high-density maraging steel were determined using a micro-spot L-DED system. Analyzing the aspect ratio and dilution relationship of beads defined a processing window for the ideal bead shape. Volumetric samples were then printed, and process parameters which created high-density material (>99%) were selected. Compared to larger laser diameters, the 0.3 mm laser significantly influences bead formation characteristics, weld pool stability, part shape accuracy, and internal defects. Internal defects such as pores and titanium oxide inclusions reduce tensile strength and ductility for this material. Variations in grain structures found throughout the deposited material were correlated with changes in cooling rate during L-DED processing. The developed L-DED repair process for the damaged HPDC insert removes the heat-checking crack and deposits material with micro-spot L-DED, restoring the component’s geometry. After the repair process, external and internal analyses provide valuable insights into the geometrical accuracy of the repair region, the microstructural and microhardness differences iii between the softer deposited material, the harder base material, and the traits of the heat-affected zones. This work showed that micro-spot L-DED is feasible for repair, but it also showed key areas in the repair methodology that require improvement before industrial implementation. Additionally, the unique response of weld beads and cuboid samples to process parameters warrants further investigation, focusing on the optimization of gas flow and powder feed rate settings for micro-spot L-DED systems. iv Lay Summary Metal additive manufacturing (AM), or 3D printing, provides new geometry and material design possibilities for any application. Within Canada, this technology is expanding into the tooling industry. In this industry, AM technology is beneficial as it can build complex components from hard-to-machine alloys (e.g., high-strength maraging steel). These alloys are used in expensive casting moulds that break after a certain number of castings. After the moulds break, they can be repaired with AM technology to lower operating costs and reduce metal waste. This thesis investigates the printing behaviour of high-strength maraging steel with a novel micro-spot laser-directed energy deposition (L-DED) system. This helps the development of a micro-spot L-DED repair process to restore a maraging steel mould to its original geometry. The thesis lays the groundwork for micro-spot L-DED printing and optimizes the process of repairing high-pressure die-casting inserts made of high-strength steel. v Preface The research conducted in this thesis was funded in part by the Natural Science and Engineering Research Council (NSERC) of Canada by way of the Canadian Graduate Scholarship Masters Award Okanagan and Canadian Graduate Scholarship Indigenous Award (#24897) and the NSERC New Frontiers in Research Fund Exploration Grant (NFRFE-2021-01025). This is a manuscript-based thesis that includes work that is either in preparation, under review or accepted for journal publication: Chapter 2 – Adapted from C. Paul, D. Nath, M.J. Benoit, L. Tobber (2024), “Review of Opportunities and Challenges for Additive Manufacturing of Steels in the Construction Industry,” Advances in Civil Engineering Materials (published). Chapter 3 – Adapted from C. Paul, J. Vincic, A. Wall, A. Jones, L. Tobber, M. Benoit (2024), “Investigation of Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging Steel Towards Repair Applications”, The International Journal of Advanced Manufacturing (under review). Chapter 4 – Adapted from C. Paul, A. Jones, J. Vincic, N. Zaker, B. S. Amirkhiz, R. Mackay, G. Byczynski, A. Elsayed, L. Tobber, M. Benoit (2024), “Laser Directed Energy Deposition Repair of Maraging Steel High-Pressure Die Cast Tooling Produced by Laser Powder Bed Fusion,” (in preparation, to be submitted to Additive Manufacturing). The use of generative AI tools in the work described here was solely for enhancing the readability and grammar of the written work. For all of the manuscripts, in any of the forms, which are included in this thesis, the author of this thesis is the lead author. The lead author led the design and performance of experiments, the formal analysis of results, the discussion of results, the vi writing of the first manuscript draft, and subsequent revisions of the manuscripts’ drafts. For these manuscripts, there are co-authors that also contributed are: - Dr. Michael Benoit: Supervision of all work performed – both formally presented in the thesis and the additional contributions –, funding acquisition, guidance on research methodology, and revision of written drafts. - Dr. Lisa Tobber: Supervision of all work performed – both formally presented in the thesis and the additional contributions –, funding acquisition, guidance on research methodology, and revision of written drafts. - Dr. Josip Vincic: Supported experimental design, performance of experiments, analysis of results, and revised written drafts. Particular involvement in the research presented in the manuscript that makes up Chapter 3. Aided in manuscript revision of Chapter 4 and collection of literature presented in the review in Chapter 2. - Dr. Alistair Jones: Supported design of repair process, analysis of results, and revised written drafts. Particular involvement in the research methodology and design of repair process presented in Chapter 4. Aided in manuscript revision of Chapter 3. - Ms. Dikshita Nath: Support for literature review and writing of the manuscript accepted at ASTM’s Advances in Civil Engineering Materials journal. Part of this manuscript is included in Chapter 1 with permission. - Robert Mackay and Glenn Byczynski with Nemak – for their technical knowledge of the HPDC insert and contributing the HPDC insert to the repair project presented in Chapter 4. Additionally, revising the manuscript that makes up Chapter 4. vii - Dr. Abdallah Elsayed: revisions of the manuscript presented in Chapter 4 with technical knowledge on cyclic thermal treatments of die casting tools to aid in the description of the damage mechanism the HPDC insert experienced. - Mr. Andrew Wall: Supported L-DED simulation procedures with Simufact Welding software utilized in Chapter 3. Additional Contributions: In addition to the main manuscripts listed above, I also contributed to research manuscripts and conference presentations (oral or poster) below: - C. Paul, L. Tobber, M. Benoit (2024), “Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging Steel”, Canadian Material Science Conference (CMSC) – 2024 (conference presentation, presenter). - A. Jones, J. Faul, C. Paul, C. Johnston, M. Benoit (2024), “3DCastleBenchy: A process-independent benchmark for additive manufacturing”, Rapid Prototyping Journal (under review, co-author). - M. K. Keshavarz, C. Paul, M. Benoit, M. Vlasea (2023), “Tailoring Microstructure and Resulting Mechanical Properties of Maraging Steel Fabricated by Laser Powder Bed Fusion”, Holistic Innovation in Additive Manufacturing Conference (HI-AM) – 2023 (poster presentation, did not present). - C. Paul, D. Nath, M. Benoit, L. Tobber (2023), “Two-Eyed Seeing in Engineering”, American Indian Science and Engineering Society (AISES) in Canada Conference – 2023 (conference presentation, co-presenter) viii Table of Contents Abstract ........................................................................................................................................... ii Preface............................................................................................................................................. v Table of Contents ......................................................................................................................... viii List of Tables ................................................................................................................................. xi List of Figures ............................................................................................................................... xii List of Abbreviations ................................................................................................................. xviii Acknowledgements ....................................................................................................................... xx Dedication ................................................................................................................................... xxii Chapter 1 : Introduction ............................................................................................................... 1 1.1. Background ......................................................................................................... 1 1.2. Goals and Objectives .......................................................................................... 5 1.3. Thesis Organization ............................................................................................ 6 Chapter 2 : Literature Review ..................................................................................................... 8 2.1. 18Ni-300 Maraging steel .................................................................................... 8 2.2. Important L-DED Parameters ........................................................................... 10 2.3. L-DED of 18Ni-300 Maraging Steel ................................................................ 12 2.4. L-DED Repair of 18Ni-300 Maraging Steel Components ............................... 16 2.5. Literature Summary .......................................................................................... 19 Chapter 3 : Investigation of Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging Steel Towards Repair Applications .............................................................................. 21 3.1. Introduction ....................................................................................................... 21 3.2. Methodology ..................................................................................................... 21 ix 3.2.1. Materials ...................................................................................................... 21 3.2.2. Experimental Procedure ............................................................................... 23 3.2.3. Microstructural Characterization and Analysis ........................................... 25 3.2.4. Finite Element Simulations .......................................................................... 26 3.2.5. Mechanical Testing ...................................................................................... 27 3.3. Results and Discussion ..................................................................................... 27 3.3.1. Single Track Experiments ............................................................................ 27 3.3.2. Cuboid Sample Fabrication.......................................................................... 35 3.3.3. Microstructural Characterization ................................................................. 39 3.3.4. Mechanical Properties .................................................................................. 44 3.4. Summary ........................................................................................................... 48 Chapter 4 : Laser Directed Energy Deposition Repair of Maraging Steel High Pressure Die Cast Tooling Produced by Laser Powder Bed Fusion .......................................................................... 51 4.1. Introduction ....................................................................................................... 51 4.2. Methodology ..................................................................................................... 52 4.2.1. Materials ...................................................................................................... 52 4.2.2. Repair Process .............................................................................................. 53 4.2.3. Microstructure analysis ................................................................................ 56 4.2.4. Mechanical Property Assessment ................................................................ 57 4.3. Results and Discussion ..................................................................................... 57 4.3.1. External Assessment of Repair .................................................................... 57 x 4.3.2. Internal Assessment of Defects .................................................................... 59 4.3.3. Microstructure Assessment of BM and FZ .................................................. 64 4.3.4. Microstructure Assessment of HAZ ............................................................ 69 4.3.5. Mechanical Properties .................................................................................. 72 4.4. Summary ........................................................................................................... 74 Chapter 5 : Extended Results and Discussion ........................................................................... 77 5.1. L-DED Weld Pool Stability .............................................................................. 77 5.2. Assessing CG and SG Settings for Micro-spot L-DED .................................... 80 5.3. Areas to Improve HPDC Insert Repair Process ................................................ 86 Chapter 6 : Summary, Conclusion, and Future Work ............................................................... 88 6.1. Summary and Conclusion ................................................................................. 88 6.2. Future Work ...................................................................................................... 92 References ..................................................................................................................................... 95 xi List of Tables Table 1: Technical, economic, and environmental benefits and challenges for fusion-based AM technology. ...................................................................................................................................... 4 Table 2: Summary of essential powder L-DED printing parameters and their main effects for single beads and multi-layer printing. ..................................................................................................... 12 Table 3: Summary of repair of steel components involving high-Ni maraging steel. .................. 19 Table 4: Nominal compositions of the 18Ni-300 powder and 304 stainless steel substrate (in wt%)........................................................................................................................................................ 22 Table 5: Printing parameters for ST experiments. ........................................................................ 24 Table 6: Parameter sets for cuboid printing. The powder feed rate is 5 g/min. The target height was 15 mm. .......................................................................................................................................... 35 Table 7: Summary of tensile properties for as-printed L-DED 18Ni-300 (n = 5). ....................... 46 Table 8: Nominal chemical compositions of powders used to print insert and for repair materials in wt%. .......................................................................................................................................... 53 Table 9: L-DED printing parameters for the repair process. ........................................................ 56 Table 10: HAZ1 and HAZ2 width data summary. The error represents the standard deviation from 15 measurements in each HAZ. .................................................................................................... 71 Table 11: Summary of Vickers hardness data. The error represents the standard deviation from 141 indents for the FZ and 243 indents from the BM, respectively. ............................................ 73 Table 12: The five parameter sets used to experiment with different CG and SG flow settings.. 82 xii List of Figures Figure 1: a) Schematics of wire arc additive manufacturing (WAAM), L-DED, and L-PBF technologies (left to right respectively) and b) weld pool schematics of the same technologies. The scan direction (SD) is indicated. Reproduced with permission from1. ........................................... 2 Figure 2: Fe – Ni phase diagram according to atomic%, adapted with permission from [36]. ...... 9 Figure 3: Schematic of single-bead a) and multi-bead b) characteristics. Images adapted with permission from [42] and [40], respectively. Criteria for avoiding lack of fusion porosity, as set by [40], provided by equations in b), H = single bead height. ........................................................... 11 Figure 4: Images (top) and schematics (bottom) of interaction areas according to laser beam diameter, S = 0.75 mm, M = 1.5 mm, and L = 3 mm. Image adapted with permission from [47]........................................................................................................................................................ 14 Figure 5: The different stages of an L-DED repair process for a hot forging die made of H13 steel: a) defect identification (wear regions indicated), b) deposition of repair material with L-DED, and c) post-repair machining to achieve surface finish. Images adapted with permissions from [69].18 Figure 6: a) SEM secondary electron micrograph of as-received 18Ni-300 powder, b) SEM backscatter electron image of powder cross-section demonstrating internal porosities (i) and particle agglomeration (ii). ............................................................................................................ 22 Figure 7: a) The LAWS250 system with labelled components: 1 – powder feeder system, 2 – laser source, 3 – water chiller system, 4 – welding chamber, 5 – LAWS programming monitor, and b) weld head inside the welding chamber: 1 – 1 of 4 powder lines, 2 – shielding cup, 3 – nozzle (contains the laser opening and coaxial powder nozzles), 4 – water coolant line, 5 – protective lens window. ......................................................................................................................................... 24 xiii Figure 8: AR and D main effects plots with respect to a) powder feed rate, b) scan speed, and c) laser power. ................................................................................................................................... 29 Figure 9: a) AR vs. D for all STs experiments with acceptable parameter sets outlined in the red dashed box, b) mass density vs. energy density for all ST experiments (beads that meet shape criteria in blue). c) cross sections of the STs inside and outside the AR and D window ranges: outside of process window - i) P=275 W, v=600 mm/min ii) P=425 W, v=1400 mm/min, inside of process window - iii) P=325 W, v=1000 mm/min iv) P=350 W, v=850 mm/min v) P=375 W, v=850 mm/min and vi) P=375 W, v=700 mm/min. The scale bar below in c) applies to all ST cross-section images. .................................................................................................................... 31 Figure 10: a) and b) bead width, c) and d) bead height, and e) and f) bead depth as a function of energy and mass density. The spot size was 2.5 mm for green ([39]) and 2.55 mm or 2.86 mm for yellow ([38]). ................................................................................................................................ 33 Figure 11: AR (a and b) and D (c and d) of weld beads versus energy density (a and c) and mass density (b and d) compared to other studies with larger laser spot diameters. The spot size was 2.5 mm for green ([39]) and 2.55 mm or 2.86 mm for yellow ([38]). ................................................ 34 Figure 12: Cuboids printed with parameter sets 1 to 8 in ............................................................. 36 Figure 13: Porosity analysis of cuboids, a) porosity distribution for all 8 samples, b) individual pore areas at different heights within set 1, set 2, and set 7, and c) pore morphologies at the different heights within set 1, set 7, and set 2, respectively. ....................................................................... 37 Figure 14: a) SEM-BSE image of the etched microstructure for layers with the two different scanning directions, b) SEM-SE images of cells and EDS map at the top – T, c) schematic showing locations in cuboid, and d) SEM-BSE images of cells and EDS map at the bottom – B of the cuboid. ........................................................................................................................................... 39 xiv Figure 15: a) – c) and d) – f) are the grain boundaries, IPF-Z maps, and IPF-Y maps of large and small area EBSD maps, respectively, from the middle to the cuboid samples. The black and white arrows in b) indicate PAGs that contain two distinct orientations of martensite blocks. BD = build direction, SD = scan direction. ...................................................................................................... 41 Figure 16: a, c) Phase and b, d) KAM maps at the a, b) top – T and c, d) bottom – B (according to location schematic in Figure 14c). ................................................................................................ 43 Figure 17: Temperature histories of the bottom and top during the printing of the cuboid sample for microstructural analysis. The lines corresponding to the liquidus (TL, 1438 °C) and solidus (TS 1042 °C) temperatures for the composition listed in Table 4 were determined using Thermo-Calc software. The red line represents the α → γ transition finish temperature of 820 °C, which is also a standard solutionizing temperature for this alloy [52], [107]. The martensite start (MS, 220°C) and finish (MF, 100°C) for 18Ni-300 are also included in the plot [52], [107], [108]. ................ 44 Figure 18: Vickers microhardness vs. build height for sample printed with parameter set 1 ( .... 45 Figure 19: Engineering stress-strain curves for five replicate L-DED 18Ni-300 samples in the as-printed condition. .......................................................................................................................... 46 Figure 20: SE-SEM images of fracture surfaces from sample two at a) 26x magnification and b) 2000x magnification, and sample four at c) 26x magnification and d) 2000x magnification. The white arrow in a) is pointing to the region analyzed in Figure 21. The black arrows in b) and d) point to small features in the dimpled fracture surfaces. .............................................................. 47 Figure 21: SE-SEM image and EDS point analysis of Titanium oxide inclusion found in Test 2 fracture surface.............................................................................................................................. 48 Figure 22: Repair process methodology outlined over six stages. In Stage 1, the heat check crack was fractured to reveal the heat check crack surface (i) and the notch-sensitized fracture surface xv (ii). Nemak provided these images. Stage 2 required a multi-step machining process, starting with ball-milling (left image) and finishing with angled end-milling to create the landing B (right image) to avoid collisions during the repair process. ................................................................................ 54 Figure 23: Photograph of repair groove after machining a) and after printing b), accompanied with dimensional colormaps created from point cloud data captured via laser scanning c) and d), respectively (top-down view of repair region). The numbers for the colour map legend are in mm units. .............................................................................................................................................. 58 Figure 24: SEM-BSE image of the entire repair region, with several defects indicated, including (A) large lack-of fusion pore, (B) spherical pores, (C) irregularly shaped defects containing a dark material, (D) non-spherical pores/voids at the groove edge, and (E) pores that were prevalent at the bottom of the groove. The black dashed box highlights porosity in the BM, while the black dashed arc indicates the bottom of the repair geometry. The build direction (BD) is also indicated........................................................................................................................................................ 61 Figure 25: Array of defects identified in the repair region, including a) and b) lack-of-fusion defects at the FZ-BM interface, c) and d) gas pores at the boundary of the weld pool, and e) and f) inclusions and small circular gas pores. The black dashed lines represent the interface between the BM and FZ. ................................................................................................................................... 62 Figure 26: SEM-SE images of the Ti/Al – O inclusion a) at 1kx magnification and b) at 5kx magnification, with c) corresponding EDS maps for Ti, Al, O, and Fe. ...................................... 63 Figure 27: a) laser deposition head and repair groove before repair and b) spatter and adhered particles on groove wall after 43 sec of the repair process. The green box outlines the side wall region, showing a) the clean wall before deposition and b) the wall with spatter particles after a few layers. ..................................................................................................................................... 64 xvi Figure 28: Etched micrographs of the a) as-built L-DED 18Ni300 in the FZ (with scanning directions indicated by the black arrow and dot) with b) the solidification structure at the edge and center of the weld track. The build direction (BD) is indicated in both micrographs. ................. 65 Figure 29: (a, c) BSE images and (b, d) corresponding EDS element maps for (a, b) the FZ and (c, d) the BM. ................................................................................................................................ 66 Figure 30: As-built L-DED in FZ a) EBSD phase, b) IPF-Y with IPF heat map, and c) grain boundary maps and HT L-PBF in the BM of the HPDC insert d) EBSD phase, e) IPF-Y with IPF heat map, and f) grain boundary maps. ......................................................................................... 67 Figure 31: The HAZ interface with different regions labelled a) with b) the hard transition between the HAZ1 and FZ. The HAZ width measurements were measured horizontally according to the “w” label. The white arrows indicate bright particles of reverted austenite. The build direction (BD) is indicated in both micrographs. ......................................................................................... 70 Figure 32: a) EBSD phase map, b) IPF-Y map and c) grain boundary map at the L-PBF to L-DED interface......................................................................................................................................... 72 Figure 33: Vickers hardness data (HV0.5): map around the repair region with the black and white dashed lines enclosing the HAZ, i – indent performed on a defect during the indentation process, scale bar representing 50 μm. ....................................................................................................... 73 Figure 34: Weld pool stability categories for STs that meet shape criteria a) and STs that do not meet the shape criteria b). Weld pool stability assessed based on inspection of weld bead surface........................................................................................................................................................ 78 Figure 35: Image of weld beads in terms of weld pool stability a) “Good”, b) “Satisfactory,” and c) “Poor” that correspond to labels i, ii, and iii, respectively, in Figure 34 a) above. Scan direction (SD) indicated in a) applies to all images. .................................................................................... 79 xvii Figure 36: Micrograph of an ST cross-section of new potential optimum parameters if weld pool stability is considered. P=200 W, v = 1000 mm/min, and 𝑓𝑝 = 3 g/min. The AR and D for this parameter set is provided. ............................................................................................................. 80 Figure 37: Images taken from the weld monitoring camera in the LAWS250 during printing to assess spatter according to parameter sets 2, 3, and 5 at CG/SG=0.3, CG/SG=1, and CG/SG=3.83 Figure 38: Weld bead surface images for the five parameter sets at CG/SG=0.3, CG/SG=1, and CG/SG=3. The scale bar at the bottom of Set 5, CG/SG=3, image represents a distance of 700 μm and applies to all images. .............................................................................................................. 84 Figure 39: Cross-sectional bead micrographs according to parameter set at CG/SG=0.3, CG/SG=1, and CG/SG=3. The scale bar at the bottom of all Set5, CG/SG=3, image represents a distance of 400 μm and applies to all images. ................................................................................................. 84 Figure 40: Bead a) height, b) width, c) depth, d) aspect ratio, and e) dilution at different CG/SG ratios for the five parameter sets. .................................................................................................. 85 xviii List of Abbreviations Al Aluminum AM Additive Manufacturing AR Aspect Ratio bcc Body centered cubic bct Body centered tetragonal BD Build direction BM Base metal BSE Back scatter electron C Carbon CAD Computer-aided design CG Carrier gas CNC Computer numerical controlled Co Cobalt D Dilution EB Electron Beam EBSD Electron backscatter diffraction EDS Energy dispersive x-ray spectroscopy El Elongation fcc Face centered cubic FZ Fusion zone GTAW Gas tungsten arc welding g Grams HAGB High angle grain boundaries HAZ Heat affected zone HPDC High pressure die casting HRC Rockwell Hardness - C scale HT Heat treatment HV Hardness Vickers IPF-Y Inverse pole figure - y direction LAWS Liburdi Automated Welding System L-DED Laser Directed Energy Deposition L-PBF Laser Powder Bed Fusion LSS Laser spot size L-MAGB Low-medium angle grain boundaries min Minute mm Millimetre Mo Molybdenum Ni Nickel OEM Original equipment manufacturer OM Optical microscopy O Oxygen PAGB Prior austenite grain boundaries PCA Point cloud acquisition xix RA Retained austenite SD Scanning direction SE Secondary electron SEM Scanning electron microscopy SG Shielding gas ST Single track Ti Titanium UTS Ultimate tensile strength WAAM Wire arc additive manufacturing W Watts YS Yield strength xx Acknowledgements First, I would like to thank my supervisors Dr. Michael Benoit and Dr. Lisa Tobber for guiding and supporting me through my master’s program. You have both taught me so much; your mentorship and encouragement helped me push through challenging times and enhance my professional, personal, and technical skills. These skills will be with me the rest of my life and for that I am thankful. I would specifically like to thank Dr. Michael Benoit for continually supporting me and my work from afar and for Dr. Lisa Tobber for “adopting” me into the ASSET Group at UBCO. To my committee members, Dr. Ian Foulds and Dr. Alexander Uhl, thank you for your technical insights and providing useful feedback that helped shape my research journey. Specifically, to Dr. Alexander Uhl, thank you for mentoring me early on as I entered the world of research during my undergraduate degree, and encouraging me to go to graduate studies. To all the members of the ASSET Group, thank you for taking me in as part of the group, your friendship and technical guidance, regardless of differing technical backgrounds, made it truly easier to show up and work. To Joel Liman and Rachel Andreychuk (and all of the staff and people at the Indigenous Programs and Services at UBCO), if not for your friendship and wise guidance, I may have never gone into graduate studies. For that, I thank you. In order of appearance, I would like to thank my close colleagues and friends I met along the way: Taha Waqar, Andrew Wall, Zahra Khodamoradi, Matt Kobel, Jordan Kozakevich, Dikshita Nath, Janelle Faul, Mitchell Barret, Kyle Lessoway, Seyedmohammad Tabaie, Josip Vincic, Alistair Jones, and Benjamin Orchard. You all provided valuable knowledge and true friendship that I will be forever grateful for. I will always fondly recall the times spent trouble shooting in the lab, discussing ideas, and spending time outside of the lab. You have all impressed lessons within me that I will always be grateful for. To all of those that I shared research space xxi with at the Innovation Precinct 1 building, thank you for the letting me borrow supplies, talk through ideas, and borrow your time. To the technicians and staff at the School of Engineering, specifically Raymond Seida and Durwin Bossy, thank you for your machining expertise and assistance to manufacture important fixtures and items along the way. To Dr. Sudip Shrestha at UBCO FiLTER, thank you for training me on the scanning electron microscope (SEM) and having the patience with a novice SEM user. To Ms. Emma Pugsley, thank you for training me, and the rest of the crew, on the LAWS250 system, which was integral to the completion of my thesis. To the Kamloops Indian Band, thank you for supporting me all throughout my post-secondary education, I can confidently say without your support I would not be where I am today. To the Syilx Okanagan people, specifically of Westbank First Nation, thank you for welcoming me here on your territory, I will always reciprocate the hospitality on my home territory in the Shuswap. Last, and certainly most important, I would like to thank my family and friends who have supported me throughout this journey. To my parents, Mary, and Frank, thank you for your loving support through school and life. You have instilled and taught me hard work and honest values that I will forever be grateful for. To my older brothers, Andrew, and Benjamin, thank you for being there and listening/sharing words of encouragement when I needed it. To my girlfriend, Jordyn, thank you for being there by my side and never failing to make me laugh. To my friends, thank you for the laughs and great memories throughout this time. I will never be able to repay you all, but I just hope to show you all how much I appreciate you. xxii Dedication I dedicate this thesis to all those I thanked above. I also want to dedicate this thesis to the seven generations of Indigenous people on Turtle Island, especially those who become engineers after me. It is a difficult path but one worth travelling down. “Do, or do not. There is no try.” -Master Yoda1 Chapter 1: Introduction 1.1. Background1 Additive manufacturing (AM), also called 3D printing, is a group of advanced manufacturing processes that build components layer by layer. AM combines automated motion systems with a form of material deposition, or formation, along the tool path. The tool path is typically defined by a ‘sliced’ 3D model, which creates a set of movement instructions in a computer numerical control language of the part to be printed [1]. In metal AM, there are two main processing techniques: solid-state and fusion-based AM. Solid-state metal AM includes technologies such as binder jetting, metal extrusion, ultrasonic AM, and cold spray AM [2]. Fusion-based AM includes technologies such as laser powder bed fusion (L-PBF), laser-directed energy deposition (L-DED), EB-melting, and arc-DED [3], [4]. The distinguishing feature of fusion-based AM is a process in which a heat source (i.e., laser, electron beam (EB), or electric arc) is used to locally melt (transitioning the metal into a liquid phase) a powder or wire feedstock, and substrate or previously printed material layer. Local melting results in the rapid solidification of a molten pool of metal as the heat source moves along the tool path. Figure 1a shows schematics of the common fusion-based AM technologies, accompanied by basic weld pool schematics (Figure 1b). It is worth noting that L-PBF differs from L-DED and wire arc additive manufacturing (WAAM) technologies by selectively melting a specific pattern in one layer of powder before another layer is spread for melting, which is repeated until the tool path is finished. Conversely, L-DED and WAAM technologies project powder or wire feedstocks into the focused region of the 1Part of the contents of this chapter have been adapted with permission from ASTM from the original publication: C. Paul, D. Nath, M.J. Benoit, L. Tobber (2024), “Review of Opportunities and Challenges for Additive Manufacturing of Steels in the Construction Industry”, Advances in Civil Engineering Materials (published). 2 respective energy source (i.e., laser or arc) to be melted as the heat source moves according to the tool path; for these technologies, this process is repeated until the end of the tool path. Figure 1: a) Schematics of wire arc additive manufacturing (WAAM), L-DED, and L-PBF technologies (left to right respectively) and b) weld pool schematics of the same technologies. The scan direction (SD) is indicated. Reproduced with permission from1. Fusion-based AM technologies present opportunities for high-value industries such as the manufacturing, aerospace, medical, and automotive industries, where the high cost can be justified [5]. In these industries, components are often made of high-cost metal alloys such as titanium, aluminum, and high-strength steels (e.g., H13, W360, and 300M), which fusion-based AM technologies are capable of processing with relative ease [6], [7], [8], [9]. Depending on the AM technology used, different applications and benefits are possible. For instance, L-PBF can produce parts that have complex and unique geometries (e.g., internal cooling or lubrication channels) with high geometric resolution, with minimal manufacturing effort, but the part size is restricted and remains relatively small [10], [11]. Conversely, WAAM is capable of creating larger components compared to L-DED and L-PBF in relatively fast times, but the geometric resolution of printed features is poor [12], [13]. Lastly, L-DED offers the ability to deposit material on pre-existing components, making repair and refurbishment of existing components possible [14], [15]. 3 Fusion-based AM technologies offer new design opportunities that are difficult to produce by conventional manufacturing methods, such as functionally graded compositions and properties and internal features [4]. Fusion-based AM also offers potential environmental and economic benefits through ‘print on demand,’ near-net shaping, repair, part consolidation capabilities, and potential use of recycled metals [16], [17], [18]. The implication of fusion-based AM technologies in high-cost industries, however, is not without challenge. Due to its niche nature and incomplete or unfamiliar global standards, regardless of relative maturity in the technology’s fundamentals, AM technologies have faced hesitant adoption in industries for a number of reasons. From an economic and technical standpoint, fusion-based AM has high up-front costs for machines and feedstock materials, and process and material reliability concerns [5]. Although, in some instances, AM-produced metals have been known to exceed their wrought counterparts in mechanical properties, concerns of anisotropy, variability of mechanical properties, and consistency of AM parts due to defects such as pores, surface roughness, and cracks still remain [19]. From an environmental standpoint, evidence has suggested that fusion-based AM processes require more energy than conventional techniques. Additionally, the lack of standard methods to evaluate and measure environmental impacts results in limited evidence supporting AM's environmental benefits [20], [21]. However, the current lack of evidence to support the positive environmental aspects of AM in general only means that there is space for research in this area [22]. Moreover, in adopting AM, engineers and technicians will need to learn new design skills that exploit AM technologies (i.e., design for AM, DfAM) and the manufacturing niche that fusion-based AM suites should best be met (e.g., low batch size, on demand, manufacturing) [16], [23]. Table 1provides a summary of the benefits and challenges of fusion-based AM discussed above. 4 Table 1: Technical, economic, and environmental benefits and challenges for fusion-based AM technology. Perspective Benefits Challenges Technical • Complex geometries can be fabricated (e.g., internal geometries). • Improved strength. • Light-weighting of designs. • Innovative design opportunities (e.g., functional grading). • Property anisotropy. • Potential for solidification-based defects (e.g., pores, cracks, residual stresses, etc.) reducing mechanical performance. • Reduced corrosion and fatigue performance compared to wrought equivalent. Economic • Part consolidation reduces assembly time. • Reduced lead time through on-demand & on-site printing for some processes. • Ability to repair damaged and refurbish existing components using some AM processes. • Limited build volumes with controlled environments for some processes. • High capital costs. • Highly trained operators. • Low productivity rates and volumes compared to conventional. Environmental • Reduces waste by precise material deposition. • Reduce material consumption. • More energy-intensive than conventional manufacturing processes. • Evidence needed to support environmental benefits. AM technology specifically compliments the manufacturing and tool-making industry because it already typically employs high-cost materials and systems [24], [25]. In this industry, L-DED shows promise as an efficient and economical option for repairing damaged dies and moulds used in casting, forging, stamping, etc., making these processes more sustainable and economical [15]. Using L-DED to repair damaged casting dies made of hard-to-machine tool steels (e.g., maraging steel) back to service condition rather than replacing them reduces high-value metal waste, extends tool life, and lowers production costs [15], [24]. Maraging steels, such as 18Ni-300 (UNS 1.2709, X3NiCoMoTi 18-9-5), are part of an ultra-high strength class of steels, which contain a low amount of carbon (C) and a high amount of nickel (Ni). Their high strength makes them good candidates for casting dies and moulds but difficult to work with and machine, especially once they have been heat treated [26]. 5 L-DED technologies are continually being improved upon by commercial vendors to enhance performance and extend the reach this technology has across industries. The L-DED system employed in this thesis is the Liburdi automated welding system (LAWS250), a novel and L-DED system that utilizes a 0.3 mm laser spot size, enhancing the abilities of L-DED technology to repair fine detailed and thin-walled components possibilities. But with the advancement of micro-spot L-DED systems, defined in section 2.3, the L-DED knowledge base needs to expand and consider micro-spot L-DED systems because challenges arise in defining suitable process parameters to achieve high-quality material. This thesis outlines the research that is required to develop an L-DED repair process for a high-pressure die casting (HPDC) insert with a micro-spot system. Once this repair process is fully optimized, it will prevent the vast production of metal waste and save production costs. The first portion of the research herein focuses on parameter optimization and material characterization of 18Ni-300 printed with a novel micro-spot L-DED system and then practical application research to design an L-DED repair process and characterize fundamental material properties of the repair region. 1.2. Goals and Objectives The ultimate goal of this thesis is to successfully repair a damaged 18Ni-300 maraging steel HPDC insert with a novel, micro-spot, L-DED system and understand the quality of the repair. To achieve this goal, the main objectives of this thesis are: 1. Characterize the weld bead shape behaviour of 18Ni-300 maraging steel printed with a micro-spot L-DED system, compare it to typical L-DED systems with larger laser spot sizes, and provide guidelines for L-DED process parameters for micro-spot systems. 6 2. Print high relative density cuboids that are geometrically accurate and quantify microstructural and mechanical characteristics of as-built 18Ni-300 maraging steel. 3. Develop a repair process that successfully and accurately deposits material in the repair groove machined into the damaged HPDC insert (referred to as “insert” throughout the rest of the thesis). 4. Quantify the effect the repair process has on the overall insert geometry and quantify the microstructural differences between the material deposited during the repair and the material that makes up the insert. Understanding the microstructural changes in the heat-affected zone (HAZ) of the insert and mapping the micro-hardness differences across the repair region is also important to assess feasibility. 1.3. Thesis Organization This thesis is constructed from a series of manuscripts that were created throughout the research program. First, parts of Chapter 1 contain content that was originally reported in a review article published in ASTM’s Advances in Civil Engineering Materials: C. Paul, D. Nath, M. Benoit, L. Tobber (2024), “Review of Opportunities and Challenges for Additive Manufacturing of Steels in the Construction Industry”, Advances in Civil Engineering Materials (accepted). After the literature review in Chapter 2, two results, and discussion, chapters are presented. These chapters, Chapter 3 and Chapter 4, respectively, are either submitted for review and/or currently in preparation: Chapter 3 – C. Paul, J. Vincic, A. Wall, A. Jones, L. Tobber, M. Benoit (2024), “Investigation of Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging 7 Steel Towards Repair Applications”, The International Journal of Advanced Manufacturing (under review). Chapter 4 – C. Paul, A. Jones, J. Vincic, N. Zaker, B. S. Amirkhiz, R. Mackay, G. Byczynski, A. Elsayed, L. Tobber, M. Benoit (2024), “Laser Directed Energy Deposition Repair of Maraging Steel High Pressure Die Cast Tooling Produced by Laser Powder Bed Fusion”, (in preparation, to be submitted to Additive Manufacturing). Chapter 3 focuses on Objectives 1 and 2 while Chapter 4 focuses on Objectives 3 and 4. Chapter 5 is a discussion of the overall outcomes and limitations of this research, supported by additional results related to parameter selection, and adjusting gas flow parameters, for micro-spot L-DED and limitations on the repair process. Finally, Chapter 6 is a summary and conclusion of the work done during this thesis and recommendations for future work related to micro-spot L-DED printing and repair. 8 Chapter 2: Literature Review This chapter provides a review of L-DED printing with 18Ni-300 maraging steel and AM repair of 18Ni-300 maraging steel components. It begins with an introduction to 18Ni-300 maraging steels and an overview of L-DED parameters, as well as the effect they have on important printing characteristics. After this, a review of literature that considers the processing of 18Ni-300 maraging steel with L-DED, with close attention paid to laser spot size (LSS), is presented. Lastly, the repair of steel components using 18Ni-300 maraging steel and various welding/repair technologies are reviewed to outline how the current repair project fits into the landscape of repair literature. 2.1. 18Ni-300 Maraging steel 18Ni-300 maraging steels nominally contain ≤0.03 wt% C and 17 – 19 wt% Ni, with alloying constituents such as cobalt (Co), molybdenum (Mo), titanium (Ti), and aluminum (Al). The term “maraging” stems from the martensitic (“mar-”) and age hardening (“aging”) nature of this alloy, combining them to make the word. The high content of alloying elements, particularly Co and Mo, means that maraging steel carries a higher cost [26], [27]. Their high strength and toughness make them widely used in high-value components such as aircraft components (e.g., landing gear, helicopter undercarriages, etc.) and armour, where their high cost (compared to more economical steels) can be justified [28], [29], [30]. In the manufacturing industry, maraging steels find applications in cutting tools, stamps, dies, and moulds [25]. The alloy’s compositional design results in strengths ranging from 1500 – 2300 MPa with appreciable ductility in the range of 10 – 18% elongation-to-failure [28]. The mechanical properties of this alloy are highly dependent on the processing history (i.e., solutionizing vs austenizing vs direct aging, aging, type of AM process used, etc.). The 9 uncharacteristic ductility of this high-strength steel, due to the lack of brittle carbides, gives it high toughness. This grade of maraging steel has a martensitic structure and achieves high-strength properties from an aging process that causes the precipitation of strengthening intermetallics. Figure 2 shows that at low temperatures and low amounts of Ni, a Fe – Ni alloy would have a ferritic (α-Fe) structure, which is austenitic (γ-Fe) at elevated temperatures. However, with increased amounts of Ni in this alloy and quenching/rapid cooling during heat treatment (HT), martensite formation is favoured at room temperature instead [26], [28]. One common misconception of low C, high Ni, maraging steels is that they possess a body-centred-tetragonal (bct) crystal structure when, in fact, they possess a softer body-centered-cubic (bcc) crystal structure [31], [32], [33]. This results in a softer martensitic microstructure compared to conventional martensitic C steels, which is beneficial for increasing ductility without tempering. The main strengthening precipitate phases that form as a result of aging are Ni3(Mo, Ti) and Ni3Al, appearing in high volumes due to the alloy-rich composition of the steel. Fe2Mo (also called Laves phases) can also form, but these can be detrimental to the ductility of the material while increasing the hardness and strength [28], [34], [35]. Figure 2: Fe – Ni phase diagram according to atomic%, adapted with permission from [36]. 10 Traditional HT of maraging steels to reach peak condition involves an austenizing step (at ~850 °C for 1 – 2 hours), followed by quenching or air cooling, and then an aging step (480 – 500 °C for several hours). The austenitizing step homogenizes the Fe-Ni matrix, while the aging step creates harder precipitates in the matrix that limit dislocation movement [28]. Over-aging (aging at 600 – 650 °C) causes enhanced austenite reversion, softening the matrix [37]. It is worth noting that in this thesis, the as-built condition (no austenizing or aging HT) of 18Ni-300 is of particular interest because HT is not required to determine the quality of printed material, and non-heat-treated material is more representative of the material condition present after a continuous L-DED repair process. Lastly, the low C in 18Ni-300 maraging steel improves its weldability characteristics and makes it a good candidate as an L-DED feedstock material [6], [25]. 2.2. Important L-DED Parameters L-DED printing of 18Ni-300 maraging steel requires the selection and understanding of various variables. The main input parameters that need to be determined are laser power, scan speed, powder feed rate (wire feed rate if applicable), layer thickness (specific to multi-layer prints), hatch spacing or overlap between adjacent tracks (specific to cladding or multi-layer prints), carrier and shielding gas flow rates, standoff distance (often defined by OEM), and laser diameter (often defined OEM). These parameters have an array of effects on weld bead dimensions, weld bead aspect ratio (AR) and dilution (D), and relative density of cuboids [38], [39], [40]. In particular, laser power, scan speed, powder feed rate, and gas flow rates are critical for proper bead formation, which is the basis for fusion-based AM processes [41]. Figure 3 a and b, respectively, show single-bead and multi-bead characteristics, such as bead width (w in Figure 3a) and lack of fusion porosity (labelled “porosity” in Figure 3b). AR is defined as the ratio of bead width to bead height above the substrate (w and h, respectively, in Figure 3a), and D is defined as 11 the ratio of overall bead over height to bead remelting depth (b in Figure 3a); equations are provided in section 3.2. Relative density is the percentage of voids, pores, or defects present in a build subtracted from one hundred (i.e., a lower relative density means greater porosity). The lack of fusion pore shown in Figure 3b happens as a result of improper selection overlap height (OH), analogous to layer thickness, and overlap depth (OD), analogous to hatch spacing, in Table 2. Figure 3: Schematic of single-bead a) and multi-bead b) characteristics. Images adapted with permission from [42] and [40], respectively. Criteria for avoiding lack of fusion porosity, as set by [40], provided by equations in b), H = single bead height. A summary of the essential L-DED parameters is presented in Table 2 with their corresponding main effects. Weld bead characteristics and relative density are essential fundamentals for fusion-based AM and the research presented in this thesis. These variables also have significant interaction effects on the outcome variables [19], [39], [41], [43], making the parameter optimization process difficult. In addition to the L-DED parameters presented in Table 2 affecting weld bead shape and relative density, they also impact the as-built microstructure and mechanical properties. The reviews by Bajaj et al. [6] and Fayazfar et al. [19] provide extensive discussions of the dependency of microstructural and mechanical characteristics on AM parameters and post-processing HT. 12 Table 2: Summary of essential powder L-DED printing parameters and their main effects for single beads and multi-layer printing. L-DED Parameter Main Effects References Laser Power Relative porosity and D. [19], [39],[38] Scan Speed Relative porosity, bead height and width, and AR [19], [39], [38] Feed rate Relative porosity, AR, and D [19], [39], [38] Layer thickness Lack-of-fusion voids. [44], [40] Hatch Spacing Lack of fusion voids and layer thickness. [40] Carrier Gas Delivery of powder to the molten pool and spatter. [45] Shielding Gas Oxidation of molten metal and build shape stability. [45], [46] Standoff distance Powder utilization efficiency and bead quality. [45] Spot Size Weld bead size, D, and powder utilization efficiency. [47], [48] 2.3. L-DED of 18Ni-300 Maraging Steel2 Several studies have explored L-DED with 18Ni-300 using a range of laser spot sizes (LSSs) – defined as the diameter of the laser beam at the working distance – with varying goals. This section introduces the progress and trends of L-DED technology with 18Ni-300 maraging steel. Chen et al. [49] (1.2 mm LSS) found that high-density parts could be achieved with an energy density of 250 J/mm2. The decreasing temperature gradient the molten pool experiences increases cell/grain size and reduces hardness along the building height. Aleksandr et al. [50] (0.96 mm laser spot) found that low energy deposition (<64.9 J/mm2) is favourable for ideal bead aspect ratio (AR, 3 – 5) and that highly dense cubes with low compressive residual stresses, a 34.7 J/mm2 energy density, 30% overlap, and bidirectional scanning strategy is recommended. Félix-Martínez et al. [38] (2.55 mm and 2.86 mm LSS) and Asala et al. [39] (2.5 mm LSS) studied the optimization and modelling of L-DED parameters for bead AR and dilution (D) for 18Ni-300 and 18Ni-350 maraging steels. While both studies created AR and D models based on extensive statistical 2Part of the contents of this chapter have been adapted with permission from: C. Paul, J. Vincic, A. Wall, A. Jones, L. Tobber, M. Benoit (2024), “Investigation of Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging Steel Towards Repair Applications”, The International Journal of Advanced Manufacturing (Under review). 13 analysis, they did not account for all parameters. For example, in Félix-Martínez et al., the powder feed rate is held constant (at 14 g/min), but spot size is included in models, and Asala et al., the spot size is held steady, but the powder feed rate is included. From these models, laser power is identified as the dominant parameter over bead D, scan speed is dominant over AR, and powder feed rate is ranked third for both [38], [39]. L-DED of 18Ni-300 has also been investigated extensively from a mechanical performance and microstructural point of view. Gil et al. [51] (2.5 mm LSS) reported that as-built L-DED 18Ni-300 had a yield strength (YS), ultimate tensile strength (UTS), and elongation at fracture (El) range from 926 – 988 MPa, 1189 – 1209 MPa, and 9 – 18%, respectively. Additionally, they observed minimal anisotropic behaviour in strength values, but more was noticed in elongation values [51]. Jiang et al. [52] (3 mm LSS) investigated the effect of different heat treatments on the hardness, wear resistance, and corrosive properties of 18Ni-300 processed with L-DED. In comparison to the as-deposited state, a direct aging treatment led to a 68% increase in hardness (380 HV versus 638 HV) and a 118% increase in wear resistance (3.81E-6 mm3/Nm versus 1.75E-6 mm3/Nm) but inferior corrosive performance. To increase corrosion performance, a solutionizing and aging heat treatment is needed to increase the amount of reverted austenite in the matrix [52]. From a more microstructural perspective, Jagle et al. [53] (1.7 mm LSS) studied the differences in microstructural features between 18Ni-300 processed via L-DED, L-PBF, and conventional methods. They found that L-DED and L-PBF had microsegregation in the as-built state and retained austenite, which was not present in the conventionally produced material [53]. In peak-aged states, the hardness of conventionally produced 18Ni-300 was not surpassed, but the L-DED and L-PBF 18Ni-300 due to an increased presence of retained and reverted austenite even after heat treatment [53]. Studies by Emadinia et al. (2.5 mm LSS) [54], Zhu et al. (LSS not available) 14 [55], Li et al. (1.2 mm LSS), and Truong et al. (3 mm LSS) also explored the L-DED processing of 18Ni-300. Most recently, Felicioni et al. (2 mm LSS) [56] examined the effect that shielding environment (Air or argon) and laser power had on the microstructural aspects of deposited 18Ni-300 maraging steel. They found that the shielding environment affects heat transfer mechanisms, affecting the quantity of retained austenite present in the as-built condition and that HT post-processing helps to combat the effects of variable cooling experienced by the material during processing, which leads to anisotropic/heterogeneous material properties [56]. The LSS effect was studied in part by [38] with 18Ni-300, but independent studies on changing LSS have highlighted its significance. LSS affects several printing and material attributes (e.g., mechanical properties, microstructure, process build stability, and part quality) by changing the size of the interaction area between the laser beam and powder streams [45], [47], [48]. Figure 4 provides images and schematics showing the change in size of the interaction volume at three different laser beam diameters (S=0.75 mm, M=1.5 mm, and L=3.0 mm). It is worth noting that in the Figure 4 images, no scale bar is provided, making the assessment of the interaction volume size qualitative. Figure 4: Images (top) and schematics (bottom) of interaction areas according to laser beam diameter, S = 0.75 mm, M = 1.5 mm, and L = 3 mm. Image adapted with permission from [47]. 15 Without a stationary layer of powder (such as in L-PBF), LSS plays a critical role in powder utilization efficiency for L-DED, influencing weld pool stability and quality of the weld bead. Studies have found that larger LSSs have better utilization efficiency compared to small LSSs [45], [47], [48]. LSS can also govern part porosity in builds. Large LSSs were found to result in greater porosity due to increased volume for gas-entrapped pores and decreased energy density, which also increased the lack of fusion pores [48]. With a small LSS, area energy density increases, potentially nearing a keyhole operation mode, which can lead to more keyhole porosities forming [57], [58]. Small LSSs allow for the fabrication of parts with greater geometric resolution and complexity compared to a larger LSS [45], [47]. Moreover, smaller LSSs can result in superior mechanical properties. For example, superior mechanical strengths (YS = 477 MPa and UTS = 662 MPa) of austenitic stainless steel were achieved with a 0.75 mm laser spot compared to a 3 mm spot size (YS = 350 MPa and UTS = 639 MPa) [47]. Similarly, Wołosz et al. [48] obtained a hardness of 819 HV for H13 tool steel with a 1.63 mm LSS, compared to 795 HV when using a 1.85 mm LSS. L-DED systems often use a focussed LSS ranging from 1 mm to 3 mm when studying 18Ni-300 [59], [44], [50]. However, one study did use a 0.6 mm LSS to investigate L-DED printing with C-bearing maraging steel, not 18Ni-300 [40]. Parameter optimization and mechanical/microstructural investigation of 18Ni-300 with L-DED has not accounted for micro-spot (LSS<0.6 mm) due to the current configurations and abilities of L-DED technology. Additionally, the importance of LSS seems to be overlooked, as in the notable and extensive reviews by Shamsaei et al. [60], Svetliky et al. [61], Ahn [41], and Guo et al.[62], the micro-spot L-DED is not discussed. With the growing use of L-DED to repair high-value, hard-to-machine components and the benefits of shrinking LSS for repair (e.g., minimal heat input and good 16 geometric fidelity for thin/small structures and complex parts) [15], [24], [47], there is a need to understand the deposition-property relationships with micro-spot L-DED systems. 2.4. L-DED Repair of 18Ni-300 Maraging Steel Components Apart from fabricating new parts, L-DED can deposit material onto existing components, thereby facilitating the repair or replacement of lost material. This section focuses on the review of L-DED technology in repair applications of steel components, which utilize 18Ni-300 maraging steel in some capacity. L-DED technology has demonstrated significant potential for repair applications, showing promise for various industries and materials. An early review by Briger et al. in 2011 [63] focussed on the use of laser cladding, an early version of L-DED technology, in industrial repair applications. According to the review, the uniqueness of components such as torpedo fins, drilling equipment, turbine fins, etc., made laser cladding the only economically feasible solution for repair [63]. Subsequent studies on L-DED for tooling repair have focused on two major topics: 1) identifying, optimizing, and studying essential parameters unique to L-DED repair and 2) implementing and demonstrating L-DED repair for new materials and components. In addition to typical L-DED parameters (laser power, scan speed, and feed rate), repair-specific parameters related to the shape and size of the repair region, as well as the deposition strategies used for repair. Graf et al. [14], 2012, explored the importance of the groove shape for the effectiveness of L-DED repair. They found V-shaped and U-shaped grooves, with adequate relief angles, improved metallurgical bonding between the part and the deposited material. Adequately designed grooves improved bonding by allowing the weld beads to form correctly during repair [14]. Rahman Rashid et al. [64] studied the effect of layer orientation on the mechanical properties of L-DED repaired maraging steel, showing that the choice of tool path 17 played an important role in the performance of repaired coupons. Barr et al. [65] studied the effect of repair depth and tool path strategies on the fatigue and tensile properties of M300 repaired via L-DED, finding that repair was suitable for shallow or deep repairs [65]. However, the failure of repaired tensile coupons initiated in the repair region [65]. Barr et al. [8] also demonstrated that repair material properties can match the materials of the substrate/original part by intrinsically heat treating the deposited material with inter-layer, or inter-track, delays programmed into the tool path. An L-DED process with any delay would differ from a continuous L-DED process, where the process is not interrupted. Smail et al. [66] optimized the L-DED repair volume, surface finish, and deposition strategy for stainless steel components damaged via pitting corrosion. Their study showed that the condition of the repair surface (sandblasted or as-machined) had a negligible effect on the repair success. However, a contour deposition strategy (rather than a zig-zag strategy) considerably improved the hardness homogeneity in the repair region and minimized defects [66]. L-DED repair in the tooling industry has primarily been researched and applied to tooling dies used in hot and cold forging operations that have succumbed to wear or corrosion damage [7], [67], [68], [69]. Figure 5 below shows the repair process of a hot forging steel die repaired by L-DED. These images are from the research of Devine et al. [69], and they worked to repair a worn-out hot forging steel die with a hard-facing Co-based alloy. Interestingly, after repair, the service life of the hot forging die was longer than the original design (exceeding the original tool life by 8%, a tool life extension of 108%) [69]. Most repair studies in this field focus on repairing an array of carbon-bearing tools steels such as P20, H13, D2, and W360 [7], [67], [68], [69]. C-bearing maraging steels such as 300M, steel have also been studied in repair applications [64], [65]. However, the L-DED repair of low-C high-Ni maraging steel components has not been studied in vast amounts recently. Grum and Slabe [70] studied the repair of 8Ni maraging steel die-casting 18 dies with laser surfacing technology (similar to laser cladding). Klobčar et al. [71] studied the aging performance of repaired 18 Ni maraging steel die-cast tooling, using a manual gas tungsten arc welding (GTAW) method instead of L-DED, and developed a aging model of the maraging steel welded regions. A subsequent study by Klobčar et al. [72] took this a step further and examined the thermal fatigue of repaired maraging steel components to understand its performance in once put back into service. Figure 5: The different stages of an L-DED repair process for a hot forging die made of H13 steel: a) defect identification (wear regions indicated), b) deposition of repair material with L-DED, and c) post-repair machining to achieve surface finish. Images adapted with permissions from [69]. Lately, studies that have explored the use of high Ni maraging steels in repair applications are outside the tooling industry. For example, Wang et al., in 2022 [73] developed an underwater repair process and compared the mechanical behaviour of 18Ni-300 maraging steel repaired underwater to 18Ni-300 maraging steel repaired in the air. Niu et al. [74] studied the behaviour of 18Ni-300 maraging steel deposited on a dissimilar steel alloy before repairing a broken gear tooth on the drive shaft of a marine vessel. Table 3 provides a summary of the significant repair studies involving low-C high-Ni maraging steels over the past two decades. In the tooling industry, this grade of maraging steel is used for die-casting operations of Al and injection moulding for plastics [75], [76], [77]. In Table 3 the component being repaired is not consistently a component for 19 casting and is not consistently made of 18Ni-300, leaving a gap in recent literature on the repair of 18Ni-300 HPDC inserts. Table 3: Summary of repair of steel components involving high-Ni maraging steel. Study Year Component Repair Technology Repair Material [70] 2004 8Ni maraging steel die casting die Laser surfacing 8Ni maraging steel [71], [72] 2008, 2012 18Ni maraging steel die casting mould Manual GTAW 18Ni-300 maraging steel [78] 2015 H11, H13, and H21 HPDC dies Laser-wire, direct metal deposition, laser-engineered net shaping. M250 maraging steel (18 wt% Ni) [73] 2022 18Ni maraging steel underwater plates Laser direct metal deposition 18Ni-300 maraging steel [79] 2022 H13 casting dies L-DED 18Ni-300 maraging steel [74] 2023 42CrMo steel gear L-DED 18Ni-300 maraging steel [80] 2023 S45 carbon steel structural part L-DED M789 maraging steel (10.3 wt% Ni) [81] 2023 20Mn2SiCrMo bainitic steel railway crossing noses L-DED 18Ni maraging steel [82] 2023 Steel hydraulic slider L-PBF Maraging M300 steel 2.5. Literature Summary Maraging steels such as 18Ni-300 are known for their high strength due to alloying elements such as Ti, Co, and Mo, which also raise their cost. Due to their exceptional strength and toughness, these steels are often used in moulds and dies in the manufacturing industry. The high cost and low-C content of 18Ni-300 maraging steel make it very suitable as feedstock for L-DED. Extensive investigations into the L-DED processing of 18Ni-300 maraging steel reveal that most systems employ a LSS of 1 – 3 mm. In L-DED, the LSS is typically a fixed parameter, requiring users to adjust other parameters to achieve high quality material. The processing knowledge of L-DED with LSSs < 1 mm is limited, and, to the best of the author’s knowledge, no literature exists on the use of a L-DED system with a LSS less than 0.6 mm. Micro-spot L-DED systems (e.g., 20 LSS<0.6 mm) could offer significant benefits for repairing complex components, but this potential has been unexplored. Recent studies on L-DED repairs involving 18Ni-300 has primarily focused on applications outside of moulds used in casting processes. This thesis aims to address this gap by investigating the micro-spot L-DED processing of 18Ni-300 maraging steel and developing a novel repair process for a real HPDC insert with a complex geometry. 21 Chapter 3: Investigation of Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging Steel Towards Repair Applications3 3.1. Introduction Micro-spot L-DED systems are far less common than L-DED systems with an LSS of 1 – 3 mm, even though smaller LSS L-DED systems present certain advantages (e.g., finer printing resolution and increased hardness), particularly beneficial in a repair application. The limited usage of micro-spot L-DEDs in literature presents challenges for defining suitable printing parameters for highly dense parts. This chapter presents an investigation to outline L-DED parameters for 18Ni-300 maraging steel in terms of ST AR and D, geometric accuracy of deposited material, and porosity. It is also interested in the single bead geometry behaviour compared to larger LSS and the corresponding microstructural and tensile properties associated with specific parameters in the as-built condition. Understanding these aspects, especially defining appropriate printing parameters, is critical before performing a repair process. 3.2. Methodology 3.2.1.Materials A commercially available gas atomized 18Ni-300 maraging steel powder obtained from Carpenter Additive was used in this chapter. The nominal chemical composition of the powder, as provided by the manufacturer, is presented in Table 4. The d10 and d90 values of the powder, as 3Part of the contents of this chapter have been adapted with permission from: C. Paul, J. Vincic, A. Wall, A. Jones, L. Tobber, M. Benoit (2024), “Investigation of Micro-spot Laser Directed Energy Deposition of 18Ni-300 Maraging Steel Towards Repair Applications”, The International Journal of Advanced Manufacturing (Under review). 22 reported by the supplier, were 45 μm and 106 μm, respectively. Scanning electron microscope (SEM) images of the powder particles are shown in Figure 6, in which spherical particles were observed. However, along with numerous small satellite particles, some elongated or irregular particles were observed. Cross-sectional analysis of the starting powder particles (Figure 6(b)) revealed internal porosity and agglomerated particles. All samples were printed on AISI 304 stainless steel plates with a nominal composition as reported in Table 4 and were 203 × 203 × 9.5 mm3 in size. Table 4: Nominal compositions of the 18Ni-300 powder and 304 stainless steel substrate (in wt%). Material Ni Co Mo Ti Cu Al Si Mn C Cr O N Fe Powder 18.38 9.07 4.86 0.82 0.10 0.07 0.03 0.02 0.02 0.01 0.01 0.01 bal. Substrate 11.50 0.15 1.25 - 0.50 - 0.50 1.00 0.04 20.75 - 0.05 bal. Figure 6: a) SEM secondary electron micrograph of as-received 18Ni-300 powder, b) SEM backscatter electron image of powder cross-section demonstrating internal porosities (i) and particle agglomeration (ii). 23 3.2.2. Experimental Procedure In this chapter, and for the remainder of the thesis, a Liburdi Automated Welding System (LAWS) 250 L-DED system (shown in Figure 7), was used to fabricate samples and perform the repair process from a powder feedstock. The different sub-systems of the LAWS250 are labelled in Figure 7a, and Figure 7b shows the weld head inside the LAWS250 welding chamber. The weld head is a critical component of this equipment as it provides the local shielding around the weld pool and directs the powder stream to the focused laser spot. The laser is focused down the center of the weld head and passes through collimating and protective lenses. The laser source (labelled 2 in Figure 7a) is equipped with an yttrium-aluminum-garnet (YAG) multi-mode laser capable of delivering a maximum power output of 1 kW. The laser has a nominal LSS (dl) of 0.3 mm at the focal distance that delivers a 1070 – 1080 nm wavelength of light in a multi-mode-Gaussian fashion, providing a “flat-top” energy distribution. The powder hopper attached to the LAWS250 (labelled 1 in Figure 7a) is a commercial powder feeder made by Oerlikon Metco. This hopper has a heated jacket and automatic stirring system to ensure the powder flows uniformly and at a controlled rate. From the powder hopper, a single gas line carriers the powder to the weld head, where it gets split into four separate streams. The powder is then delivered to the laser focal point through four coaxial powder nozzles. Argon gas, with a purity of 99.999%, served as the carrier gas for the powder stream at a flow rate of 3 L/min and for melt-pool shielding with a gas flow rate of 9 L/min. 24 Figure 7: a) The LAWS250 system with labelled components: 1 – powder feeder system, 2 – laser source, 3 – water chiller system, 4 – welding chamber, 5 – LAWS programming monitor, and b) weld head inside the welding chamber: 1 – 1 of 4 powder lines, 2 – shielding cup, 3 – nozzle (contains the laser opening and coaxial powder nozzles), 4 – water coolant line, 5 – protective lens window. A multi-stage approach was used to determine process parameters for producing dense bulk samples involving single track (ST), single layer, and finally, multi-layer cuboid printing. STs of length 40 mm were first deposited using various combinations of laser power (P), scan speed (v), and powder feed rate (fp), as shown in Table 5. In the discussion of the ST results, bead geometries will be discussed with regard to linear energy density (Ed) and linear mass deposition (Md), both normalized against the laser spot diameter according to the following equations: Equation 1: 𝐸𝑑 [𝐽𝑚𝑚2] =𝑃𝑣∗𝑑𝑙 Equation 2: 𝑀𝑑[𝑔𝑚𝑚2] =𝑓𝑝𝑣∗𝑑𝑙 Table 5: Printing parameters for ST experiments. Experiments Power (W) Scan Speed (mm/min) Feed Rate (g/min) Initial 200, 275, 350, 425, 500 600, 1000, 1400 3, 5, 8 Refined 300,325, 350, 375, 400 700, 850, 1000 5 The STs were cross-sectioned and prepared by grinding and polishing to achieve a mirror-like surface finish. ST printing parameter sets were selected based on meeting AR and D criteria 25 ranges of 3 to 5 and 0.5 to 0.75, respectively. A set of refined experiments (Table 5) were subsequently conducted to further analyze the initial parameter sets to define a suitable window for printing parameters. Parameter sets were selected to print multi-track single-layer depositions (clads) that were 40 mm long and 20 mm wide. Successful parameter sets were selected for bulk printing with repair applications in mind. Overlap between adjacent tracks (OV%) of 30% and 40% of the bead width was used to achieve uniform layer height and high density [40], [50]. The height of each of these eight clads was measured at five locations, and the average value was used as the layer thickness(Δz) for fabricating multi-layer cuboidal samples. The cuboid samples were produced using a simple bi-directional deposition pattern, with 90° rotations every two layers. The cuboid bulk samples had dimensions of 20 × 20 × 15 mm³ and were produced using eight selected sets of parameters. Bulk samples were also printed using the parameters resulting in the highest density samples to fabricate horizontal tensile bars explained at the end of section 3.3.2. 3.2.3. Microstructural Characterization and Analysis ST and cuboid samples were cross-sectioned and prepared by conventional metallographic grinding and polishing practices before assessing by optical microscopy using a Keyence VHX 970F digital microscope. The samples were etched in a 5-vol% nital solution to reveal the weld bead shape and microstructure. The porosity of cuboid samples was evaluated by taking 25 optical micrographs (five across the width of the sample at five locations along the building height) at 200× magnification and performing quantitative analysis with ImageJ software. A Tescan Mira 3 XMU SEM equipped with a field emission electron gun was used for SEM imaging, energy dispersive x-ray spectroscopy (EDS) to capture chemical data of the material, and electron backscatter diffraction (EBSD) to capture grain and phase information of the as-built 18Ni-300 maraging steel. The accelerating voltage used with the field emission gun 26 was 15 keV for SEM images and EDS, and 30 keV was used for EBSD. The accelerating voltage for imaging was selected to be larger than nickel's Kα principal line energy (Ni), the main alloying element in 18Ni-300 maraging steel, while trying to minimize the beam interaction volume. For EBSD, the accelerating voltage was increased to 30 keV to increase the electron beam energy. Increasing the electron beam energy increases the amount of crystal information that can be captured during diffraction by increasing the source volume of the electrons [83]. The EDS analysis was performed using an Oxford Instruments X-Max EDS detector with a minimum acquisition time of 20 min for elemental maps. EBSD analysis was performed using an Oxford Instruments Nordlys EBSD detector. A 0.1 μm step size was used for EBSD mapping. The amount of both martensite and austenite was recorded in the EBSD phase analysis, with the commonly used assumption that martensite would be indexed as a body-centred cubic (bcc) structure [53], [73], [84], [85]. After EBSD mapping, Aztec Crystal analysis software was used to clean the data and analyze the results. 3.2.4.Finite Element Simulations To further explore the spatial variation in cooling rates and resultant microstructures, finite element L-DED simulations were performed using the welding and AM simulation software Simufact Welding (Hexagon, Sweden). This allowed the temperature histories during printing to be predicted at the same heights where the microstructural analysis occurred. In the Simufact simulation, the laser power, scan speed, and laser spot diameter were all set to match the experimental setup – 375 W, 700 mm/min, and 0.3 mm, respectively. The heat source efficiency was set to 0.3, and the Gaussian parameter was set to 0 to give the laser beam a flat-top power distribution across the beam profile. Additionally, an analytical cooling rate (?̇?) equation that is based on the primary cell arm spacing (PCAS) was used to calculate cooling rates based on 27 microstructural features [86]. PCAS was calculated using the line intercept method with cells in micrographs. This equation has been used to calculate the cooling rates of rapidly solidified steels in the past by Fu et al.[87] and Mithal et al. [88] (applied to L-DED) thus, it was deemed suitable for this application. 3.2.5.Mechanical Testing Vickers hardness testing was conducted using a Wilson VH3100 Vickers hardness indenter with a 500-gram force load and a 12-second dwell time. Mapping of hardness in each of the cuboid samples was captured using an array comprised of seven columns of indentations along the build height, with each column spaced 1 mm apart and indents spaced 0.25 mm apart along the build direction. Tensile testing was performed on samples fabricated with the processing parameters that produced the lowest porosity in cuboid samples. Tensile samples were evaluated only in the horizontal orientation due to process instability with tall and slender bulk samples. The samples were machined from horizontally printed rectangular prisms with 108 × 10 × 7 mm3 dimensions. Rectangular tensile specimens were designed per ASTM E8 sub-size specimen specifications, including a gauge length, width, and thickness equal to 32 mm, 6 mm, and 3 mm, respectively. Tensile testing was performed on an Instron 5969 dual-column load frame at a constant crosshead speed of 0.48 mm/min. The strain data was captured with an advanced video extensometer (AVE) system integrated into the Instron load frame. 3.3. Results and Discussion 3.3.1.Single Track Experiments Initial ST experiments ranged in linear mass and energy density normalized against laser spot size from 0.007 to 0.04 g/mm2 (3 – 8 g/min) and 28.6 to 166.7 J/mm2, respectively. The refined ST experiments ranged in area mass density and area energy density from 0.017 to 0.024 28 g/mm2 (varying scan speeds at a fixed powder feed rate of 5 g/min) and 60 to 114.3 J/mm2, respectively. Figure 8 shows the effects of the three main process parameters studied during the experiments (i.e., P, v, fp) on the weld beads' AR and D. To derive the AR and D from the bead geometries, three characteristic features were measured from the cross-sectional images: width (w), height (h), and depth (d), as illustrated in Figure 9c-vi. AR and D were calculated using Equation 3 and Equation 4, respectively [89],[90]: Equation 3: 𝐷 = 𝑑ℎ+𝑑 Equation 4: 𝐴𝑅 = 𝑤ℎ 29 Figure 8: AR and D main effects plots with respect to a) powder feed rate, b) scan speed, and c) laser power. Linear regression and analysis of variance were used to quantify the effects of process parameters on the bead geometry. Powder feed rate (Figure 8a) was found to be negatively correlated with both AR and D and was determined to be the most significant parameter (determined with a p-value of 0.05) with a larger negative effect compared to the lesser positive effects of laser power (Figure 8b) and scan speed (Figure 8c). This relationship is also shown in Figure 9a, with measured AR and D values of all ST experiments according to the different mass flow rates. From the figure, D tends to increase with increasing AR, albeit at different rates for 30 different mass flow rates – the rate-change of D with respect to AR is lower for 3 g/min than 5 g/min and 8 g/min. An 8 g/min powder feed rate produces STs with low AR and D values as more material is deposited per unit length. Higher material feed rates feed more powder into the melt pool, creating low-dilution beads. Figure 9c-i represents the shape of STs in this region, and the contact angle (a.k.a wetting angle) between the bead and the substrate surface is >90°. Ideally, this angle would be <90° to avoid a potential lack of fusion between beads and increase the quality of the layer [90]. Conversely, a 3 g/min powder feed rate produces STs with high AR and D. With low feed rates, less powder is fed into the melt pool, which creates low bead growth and higher dilution beads. Figure 9c-ii represents the shape of the beads in this region; notice the flatness of the bead with minimal growth above the substrate. With a smaller laser spot size, a smaller interaction volume exists between the laser beam and the powder streams, which enhances the influence of the powder feed rates on AR and D. Small interaction volumes reduce the stability of weld pools and bead size due to low the powder utilization/capture efficiency [47], [48]. The powder capture efficiency also affects the build/layer growth stability of multi-layer prints [91]. With the shape behaviour of STs defined, those that met targeted ranges for AR and D were identified. In Figure 9a, the dashed-red box highlights the STs that meet bead shape criteria (18 in total) to achieve a balance between metallurgical bonding without excessive re-melting (0.5<D<0.75) and ellipsoidal beads that are beneficial for volumetric printing (3<AR<5) [50], [92], [93]. Figure 9b depicts the processing window in terms of energy and mass densities for beads inside the shape criteria (blue) and those outside the shape criteria (orange). For beads that met AR and D criteria, an average energy density and mass density of 74.5 J/mm2 and 0.018g/mm2, respectively, should be used. Inside the process window, four STs ranging in AR and D from 3.29 to 3.62 and 0.64 to 0.71, respectively (Figure 9ciii – vi) were identified as suitable candidates for 31 volumetric printing. These were selected because they met metallurgical bonding and volumetric printing requirements for repair applications and represented different energy and mass density combinations (low, medium, and high) in the optimal parameter window, according to Figure 9b. Figure 9: a) AR vs. D for all STs experiments with acceptable parameter sets outlined in the red dashed box, b) mass density vs. energy density for all ST experiments (beads that meet shape criteria in blue). c) cross sections of the STs inside and outside the AR and D window ranges: outside of process window - i) P=275 W, v=600 mm/min ii) P=425 W, v=1400 mm/min, inside of process window - iii) P=325 W, v=1000 mm/min iv) P=350 W, v=850 mm/min v) P=375 W, v=850 mm/min and vi) P=375 W, v=700 mm/min. The scale bar below in c) applies to all ST cross-section images. Figure 10 shows a comparison of bead dimensions between the ST experiments of the current study (n=60) and two other studies; one study conducted by Asala et al. [39] (green) with a 2.5 mm spot size, and the other study by Felix-Martinez et al. [38] (yellow) with a 2.55 or 2.86 mm spot size. The R2 values for the current study are shown in the bottom right-hand corners of the graphs. The bead dimensions are plotted with respect to energy density (Figure 10 a, c, and e) and mass density (Figure 10 b, d, and f). 32 Between studies, bead widths and heights have the most significant discrepancies. For example, in the current study, the bead width ranges from 400 to 1200 μm, and the bead height ranges from 100 to 500 μm. However, the larger laser spot studies have widths and heights ranging from 1200 to 3200 μm and 150 to 1500 μm, respectively. The main reason behind the difference in bead scale between LSSs lies in the different sizes of the laser-powder interaction region, best visualized in Figure 4 from the study by Kong et al [47]; Wolosz et al. [48] also provide supporting evidence for this. These studies found that the larger laser-powder interaction region, using larger LSSs, creates larger bead dimensions because the larger spot is able to capture more powder. Figure 10a shows that the beads in [39] are over twice the width of the current study and have a similar energy density range. From Figure 10c, similar bead heights can be produced by both studies at low energy densities. However, as energy density increases, the studies that used larger LSSs created taller beads. Considerable overlaps in bead depth (Figure 10e and f) are seen between studies. The R2 values for the current study indicate no strong relationship between either energy or mass density and bead width, height, or depth. One exception, though, is the relationship between bead height and mass density, where a strong positive correlation with an R2 = 0.87 exists. Ascari et al. [94] found a similar relationship between linear mass density and layer height (STs overlapped together to make a single layer), with an increase in the mass density increasing the layer height. 33 Figure 10: a) and b) bead width, c) and d) bead height, and e) and f) bead depth as a function of energy and mass density. The spot size was 2.5 mm for green ([39]) and 2.55 mm or 2.86 mm for yellow ([38]). Figure 11 compares bead AR and D between large and small LSSs. This figure shows that even though different spot sizes are used, and the bead cross sections vary in dimension, similar AR and D values are achieved. Despite small spot sizes, D seems higher than the D of large laser spot sizes. The range of D values with a 0.3 mm spot size does not go below 0.3 and has a higher average value of 0.65±0.13. This also shows that AR and D alone cannot capture differences in scale between smaller and larger LSSs. Further, the data emphasizes the importance of powder flow rates with small spot sizes. The mass density for the current study has a much higher R2 value for AR and D than energy density. The negative correlation between AR and energy and mass 34 density (albeit not strong) was also found by Aleksandr et al. [50]; they did not provide an explanation for the relationship between energy or mass density and D. Figure 11: AR (a and b) and D (c and d) of weld beads versus energy density (a and c) and mass density (b and d) compared to other studies with larger laser spot diameters. The spot size was 2.5 mm for green ([39]) and 2.55 mm or 2.86 mm for yellow ([38]). Smaller than typical LSSs present both benefits and challenges for L-DED processing. First and foremost, small LSSs create smaller beads that can be beneficial for forming accurate, small, and complex geometries [47]. Additionally, for austenitic stainless steels and H13 tool steel particularly, utilization of smaller spot sizes leads to increased yield strength and hardness due to microstructure refinement [47], [48]. Unfortunately, small beam diameters are detrimental to feedstock utilization/capture efficiency during L-DED printing, which has many negative side effects. For instance, low material capture efficiency leads to unstable weld pools that contribute to poor surface finish quality, material quality (i.e. defects as shown in 3.3.2) and productivity [45], [47]. With narrower and thinner beads that have increased D (Figure 10 and Figure 11), tool path planning and slicing strategies are affected; narrower and thinner beads mean more beads are 35 required to make a single layer and more layers are required to make a part. More beads and layers could result in more heat input during printing, potentially undesirable for repair applications. 3.3.2.Cuboid Sample Fabrication Single-layer and cuboid samples were printed using the four parameter sets highlighted in Figure 9ciii-vi, using 30% and 40% overlap between STs. Table 6 contains the complete parameter sets that were used for printing cuboids, and images of all cuboids in the as-printed state are provided in Figure 12. Table 6: Parameter sets for cuboid printing. The powder feed rate is 5 g/min. The target height was 15 mm. Set (Figure 9 image) Power (W) Scan Speed (mm/min) Energy Density (J/mm2) Mass Density (g/mm2) Overlap (%) ΔZ (mm) Final Height (mm) Porosity (%) 1 (vi) 375 700 107.14 0.0238 30 0.48 14.73 0.38 2 (vi) 375 700 107.14 0.0238 40 0.53 13.43 0.55 3 (v) 375 850 88.24 0.0196 30 0.38 11.15 0.37 4 (v) 375 850 88.24 0.0196 40 0.43 12.86 0.60 5 (iv) 350 850 82.35 0.0196 30 0.49 16.34 0.60 6 (iv) 350 850 82.35 0.0196 40 0.63 15.48 0.53 7 (iii) 325 1000 65.00 0.0167 30 0.36 12.11 0.71 8 (iii) 325 1000 65.00 0.0167 40 0.40 14.29 0.67 Cuboids 1, 5, 6, and 8 all had less than 10% geometric error with respect to the intended build height of 15 mm, with cuboid 1 achieving the closest accuracy at -1.8% error. Printing of cuboid 1 was not complete due to an interruption of the process, but the error occurred on the last layer thus, it is included in these results. Cuboids 5 and 6 are the only overbuilt conditions, with positive build height errors. Cuboids 2, 3, 4, and 7 were underbuilt with more than 10 % geometric error. Edge waviness defects were pronounced on cuboids 3 and 7, while the sides of cuboids 2, 5, and 6 had pronounced satellite defects. 36 Figure 12: Cuboids printed with parameter sets 1 to 8 in Table 6. Red arrows indicate edge waviness, and green arrows indicate satellites. The distribution of relative porosity measurements for each cuboid is shown in Figure 13a (note n = 25 images per sample). Cuboid 1 had the lowest average relative porosity at 0.38%. Cuboids 2 and 3 have locations in the cross-section with large outlying data points with relative porosity > 3.5%. A visualization of the pore size (quantified by area % in OM images) vs. the height above the substrate in cuboids 1, 2, and 7 can be seen in Figure 13b. Cuboid 1 was selected as it had the lowest porosity, cuboid 2 was selected for its large outlying data points, and cuboid 7 was selected because it had the largest standard deviations. It can be seen that the maximum pore size increases with height above the substrate. Throughout the building height, two types of defects were generally observed: small spherical gas pores (Figure 13c i) and larger irregularly shaped defects filled with a dark material (Figure 13c ii and iii). 37 Figure 13: Porosity analysis of cuboids, a) porosity distribution for all 8 samples, b) individual pore areas at different heights within set 1, set 2, and set 7, and c) pore morphologies at the different heights within set 1, set 7, and set 2, respectively. The defects that can occur during L-DED printing are both external (e.g., edge waviness, satellite defects, and non-uniform layer thicknesses [95]) and internal (e.g., inclusions, gas pores, or lack-of-fusion pores [25]). Externally, cuboids 3, 4, 7, and 8 show edge waviness along the top surfaces. Edge waviness defects are caused by excess or uneven heat build-up during printing and inconsistency in material flow rate [95]. The satellite defects on the sides of cuboids 2, 5, and 6 are point anomalies instead of edge anomalies. Satellites are caused by non-optimal powder flow rates [95]. The lower powder utilization efficiency associated with small laser spot sizes could contribute to these defects as they are associated with material flow rates. The thickness of the layers used during the printing process could also be attributed to these defects. For example. cuboids 1, 2, 5, and 6 all used layer thicknesses >0.45 mm, and satellites were seen, but edge waviness was minimal. Conversely, cuboids 3, 4, 7, and 8 all used layer thicknesses <0.45 mm and minimal satellites were observed while edge effects were observed. The layer thickness, though, is more responsible for the geometric inaccuracies of cuboids. The non-uniform layer 38 growth was hard to quantify intrinsically, but this defect manifests as deviations in part height accuracy. The final build height of the cuboids printed for this chapter in Table 6 highlights deviations from the cuboid samples’ designed height (15 mm). Unstable layer growth can lead to the creation of process-induced porosity from the mismatch in the stepping distance between layers and the actual thickness of the layer [44], [91]. For example, the overbuilding of samples (Cuboid 5 and 6 only) occurred as the working distance between the surface and the nozzle became smaller during the print. If underbuilding occurred, that working distance became larger as a result of the defocused laser spot and powder streams, leading to poorer bead geometry and, therefore, less layer growth during the printing [91]. Internally, small spherical pores were observed throughout the deposited material and can be attributed to entrapped gases or potentially keyholing (Figure 13c-i). The irregularly shaped defects filled with a dark material were typically more common in the last deposited layers. Their irregular shape could be attributed to a lack of fusion (Figure 13c-iii) but the lack of fusion phenomenon alone would not explain the dark material inside. Dark material defects in L-PBF printed 18Ni-300 maraging steel have been identified as titanium oxides in the deposited material. The formation of these oxides was attributed to the remelting of spatter produced by unstable weld pools and fluctuations in the shielding environment [47], [84]. The low shielding gas flow rate (9 L/min), local shielding of the weld pool, and spatter from unstable weld pools seen during processing, commonly associated with smaller LSSs in L-DED, explain the formation of these defects. To combat weld pool instability with micro-spot L-DED systems, attention should be directed to the powder feed rate (importance illustrated in Figure 8 and Figure 9) and shielding and carrier gas flow ratios [45] – previously not of great importance in past parameter optimization studies. 39 Based on the assessment of porosity (Figure 13), geometric accuracy (Table 6), and edge characteristics (Figure 12), the parameters used for printing cuboid 1 were selected for further microstructure analysis and mechanical testing. 3.3.3.Microstructural Characterization Figure 14a shows an SEM back-scatter electron imaging (BSE) etched microstructure of the printed material. The red and blue dashed lines define regions of orthogonal laser scanning directions (SD). Within region SD1, region i consists of a fine equiaxed cell structure while region ii, at the edge of the melt pool, consists of a more columnar and elongated cell structure that grows towards region i; the solid blue lines in the figure indicate the local cell growth directions in region ii (i.e., perpendicular to the melt pool boundary). Within region SD2, the structure comprises both columnar and equiaxed cells that grew in favour of the scanning direction, tilting them to the left in the image, as indicated by the solid red lines in the figure. Figure 14: a) SEM-BSE image of the etched microstructure for layers with the two different scanning directions, b) SEM-SE images of cells and EDS map at the top – T, c) schematic showing locations in cuboid, and d) SEM-BSE images of cells and EDS map at the bottom – B of the cuboid. 40 Figure 14b and d show secondary electron (SE) images of cells, with their corresponding elemental distribution maps, at the top and bottom locations of the cuboid, respectively. The EDS maps show microsegregation of molybdenum (Mo), aluminum (Al), and Ti at the intercellular regions due to the rapid solidification within the weld pools [96]; it should be noted that Ti concentrations appear as discrete regions/particles within the microstructure. The average cell size at the bottom of the cuboid was measured to be 3.07 μm, while at the top, it was found to be 4.95 μm. The cooling rate (?̇?) can be estimated using the average primary cell arm spacing (PCAS) measured from Figure 14 using the empirical equation: Equation 5: ?̇?[°C/s] = (80𝑃𝐶𝐴𝑆)3.03 In which 80 and 3.03 are material constants for steel [86]. At the top of the cuboid, the cooling rate is calculated to be 4,584.3°C/s. At the bottom of the samples, the cooling rate is calculated to be 19,586.3°C/s, which can be attributed to the proximity to the large substrate on which the samples were printed. The lower temperature gradient at the top of the cuboid is responsible for the lower cooling rate, leading to larger solidification cells [49]. EBSD analysis was performed at low (Figure 15a-c) and higher (Figure 15d-f) magnifications to evaluate broad patterns and details of the microstructure in the middle of the cuboid sample. Black (5° – 20°) and red (> 50°) boundaries in Figure 15 indicate low-medium angle grain boundaries (L-MAGB) and high-angle grain boundaries (HAGBs), respectively. The L-MAGBs are representative of dislocations in the matrix [52], while HAGBs in as-built L-DED 18Ni-300 have been identified as blocks of martensite laths that form within packets [32], [53], [97]. Martensite packets and blocks are within prior-austenite grain boundaries (PAGBs), indicated by white boundaries (20° to 50°). From Figure 15a, PAGBs are inclined to the right, which aligns with the local laser scanning direction. The blocks of martensite in the PAGs have specific 41 orientations. For example, the black arrow in Figure 15b indicates a PAG with blocks of martensite with <101> and <001> orientations, while the white arrow indicates a PAG containing blocks with <001> and <111> orientations. In Figure 15d to f, the as-built microstructure is not easily defined as having a martensitic lath structure but instead martensite cells/blocks that contain dislocations [98]; with an EBSD step size of 0.1 μm, the fine martensitic laths would be difficult to distinguish in these maps as the lath size in as-printed L-DED maraging steels has been reported to be ~100 nm – 500 nm [73], [99]. Typical martensitic lath structures have been reported as difficult to identify in the microstructure either because they are too small or not formed yet due to the lack of heat treatment [53]. Figure 15: a) – c) and d) – f) are the grain boundaries, IPF-Z maps, and IPF-Y maps of large and small area EBSD maps, respectively, from the middle to the cuboid samples. The black and white arrows in b) indicate PAGs that contain two distinct orientations of martensite blocks. BD = build direction, SD = scan direction. Figure 16 presents the EBSD phase and kernel average misorientation (KAM) maps from regions at the top and bottom of the cuboid sample. The dominant phase in both locations is bcc, which is generally considered to be consistent with martensite in the study of printed martensitic steels [52], [53], [100], [101]. The top and bottom locations of the cuboid have 96.1% and 90.1% 42 of martensite (bcc), respectively, and 3.9% and 9.8% of austenite (fcc), respectively. Carbon-bearing martensitic steels typically possess a body-centred tetragonal (bct) crystal lattice structure [32]. However, high-Ni martensitic steels that have minimal C have been reported to consist of a softer bcc martensite crystal lattice [31], [32], [102]. The retained austenite (fcc) shown in red in Figure 16a and c is found in the inter-granular/cellular bcc regions. The retained austenite populates these regions because they are rich in Mo, Ti, and Ni, according to EDS maps in Figure 14. Mo, Ti, and Ni stabilize austenite in these regions by reducing the martensite starting temperature, creating a more stable environment for austenite [54], [103], [104]. Lastly, the grain size, classified by equivalent circle diameter (ECD), at the top and bottom are 2.17 μm and 1.53 μm, respectively, which agrees with the cell sizes calculated previously in Figure 14. KAM maps are presented in Figure 16b and d for the bcc phase only; the white spots in the figure correspond to fcc that was removed from the maps for analysis. Local misorientations are observed throughout the microstructure in both the bottom and top regions, indicating a strained microstructure consistent with martensite. 43 Figure 16: a, c) Phase and b, d) KAM maps at the a, b) top – T and c, d) bottom – B (according to location schematic in Figure 14c). Figure 17 shows the FE simulated temperature history at the top and bottom of the cuboid sample. The data shown in Figure 17 is single node temperature history data – taken from nodes at the center of the sample’s mesh. The bottom of the cuboid, which has the most fcc present at 9.85%, has an estimated cooling rate of 16,509.5 °C/s. The top of the cuboid, which has the least FCC present at 3.91%, has an estimated cooling rate of 6,281.4 °C/s. Compared to the cooling rates calculated with Equation 5 above, these cooling rates are within the same orders of magnitude, confirming that the different locations have different cooling rates. Discrepancies in cooling rate and temperature history (i.e., re-heating) throughout build volumes result in differing microstructural behaviour. High cooling rates experienced by martensitic steel during AM can result in retained austenite in the as-built microstructure [105], as was seen in Figure 16. With a higher cooling rate at the bottom than at the top, more fcc is present because the higher cooling rate prevents austenite from completely transforming to martensite through solid-state diffusionless shear transformations [32], [74], [106]. Additionally, the bottom 44 region experiences more thermal cycles to temperatures above the ferrite-austenite transformation temperature (~820°C, red α → γ line in Figure 17), leading to the formation of reverted austenite and, hence, a greater amount of fcc at the bottom compared to the top [55]. Additionally, throughout building the sample, the bottom is held above the Ms temperature (~350 °C) until the end of the printing process. In contrast, the top layers are not held at an elevated temperature for an extended period. At both locations, the Mf temperature (100 °C) is not reached until after the printing process (printing time ~1500 s). Upon cooling below Mf, the microstructure consists of austenite and martensite. Figure 17: Temperature histories of the bottom and top during the printing of the cuboid sample for microstructural analysis. The lines corresponding to the liquidus (TL, 1438 °C) and solidus (TS 1042 °C) temperatures for the composition listed in Table 4 were determined using Thermo-Calc software. The red line represents the α → γ transition finish temperature of 820 °C, which is also a standard solutionizing temperature for this alloy [52], [107]. The martensite start (MS, 220°C) and finish (MF, 100°C) for 18Ni-300 are also included in the plot [52], [107], [108]. 3.3.4.Mechanical Properties Figure 18 shows the hardness profile along the build height of the cuboid sample. The lowest hardness value along the building height is < 200 HV0.5, observed at the interface between the deposited material and the 304 stainless steel substrate. The hardness increases through the heat-affected zone towards the substrate, reaching a value of ~240 HV0.5. There is an initial rapid 45 increase in hardness with increasing build height as the indents transition into the fine martensitic matrix resulting from L-DED [54]. However, above a build height of ~0.5 mm, the hardness stabilizes around an average value of 339.3±9.91 HV0.5, which is in agreement with prior reports for as-printed L-DED 18Ni-300 [40], [52], [54], [104], [108], [109]. One interesting phenomenon observed in L-DED is in-situ heat treatment due to cyclic reheating of the material (Figure 17) [108], [110]. However, in this chapter, the hardness mapping along the building height (Figure 18) showed no indication that hardness increases/decreases with increasing build height. A lack of hardness change along the build height indicates that intrinsic heat treatment did not occur. Additionally, the thermal profiles in Figure 17 show that during printing, the temperature is only in the aging temperature range for this alloy (400 – 520°C) for a very short period [8]; studies on isothermal solidification of 18Ni-300 have shown that an optimal temperature of 490 °C for 2 – 4 hrs is required to age it a significant amount artificially, increasing hardness up to 500 and 600 HV [52], [111], [112]. Figure 18: Vickers microhardness vs. build height for sample printed with parameter set 1 ( Table 6). Figure 19 shows the tensile stress-strain curves for five replicate samples. Table 7 summarizes the average tensile properties of the tested samples. The ultimate tensile strength 46 (UTS) and yield strength (YS) of the samples are close to the values for horizontal samples from the study by Gil et al. [51]. In their research, the YS, UTS, and El were 925.9 MPa, 1188.6 MPa, and 18.3%, respectively. Comparing these values to those in Table 7, the YS values are within one standard deviation, but the UTS and El values are further apart. Figure 19: Engineering stress-strain curves for five replicate L-DED 18Ni-300 samples in the as-printed condition. Table 7: Summary of tensile properties for as-printed L-DED 18Ni-300 (n = 5). Orientation YS (MPa) UTS (MPa) El (%) Horizontal Samples 888.0 (±65.43) 1081.2 (±38.27) 10.2 (±1.4) SE-SEM images of samples 2 and 4 fracture surface are shown in Figure 20; these samples were selected for analysis as they represent the lowest EL and lowest UTS, respectively. The fracture surfaces of both Test 2 and Test 4 were dimpled, as can be seen from Figure 20b and d, indicative of ductile failure [113]. However, pores (approximately 3 μm in diameter) and regions of potential cleavage fracture can also be seen across the surface, as indicated by black arrows in Figure 20. At lower magnification, differences between samples 2 and 4 fracture surfaces become more prevalent. Particularly, sample 2 has more large pores and defects on its surface (white arrow in Figure 20a) while the fracture surface of sample 4 contains a lower amount of these larger defects. 47 Figure 20: SE-SEM images of fracture surfaces from sample two at a) 26x magnification and b) 2000x magnification, and sample four at c) 26x magnification and d) 2000x magnification. The white arrow in a) is pointing to the region analyzed in Figure 21. The black arrows in b) and d) point to small features in the dimpled fracture surfaces. A higher magnification SE-SEM image of the defect indicated in Figure 20a is shown in Figure 21. EDS point analysis of the darker material (indicated by i) revealed a more significant local concentration of Ti and O, indicating titanium oxide inclusion [49]. This coincides with the titanium oxide inclusion identified in the defect analysis section 3.3.2. The titanium inclusion has a sharp, faceted surface, suggesting local brittle fracture. Additionally, as indicated by the white arrow, the matrix in the fracture surface immediately surrounding the titanium inclusion is a flatter and less dimpled region, indicating a potential fracture initiation location. Titanium oxide inclusions are known to be problematic for the tensile properties of 18Ni-300 and other maraging steels processed with laser-based AM techniques [39], [49], [84]. The formation mechanism of 48 these oxides has been related to a fluctuation in the inert shielding environment during the AM process due to potentially unstable weld pools causing oxidized particles to be remelted in subsequent layers [38], [84]. These inclusions have been related to premature failure, resulting in less plastic deformation of samples [39], [46], [51]. Fracture surfaces presented in [51] also had pores and inclusions but at a lower amount. Similarly, analysis of tensile and fatigue fracture surfaces of L-PBF 18Ni-300 samples by Deirmina et al. [101] indicated that fracture/failure initialization stems from these titanium oxide inclusions. Thus, the suspected cause for inferior tensile performance (compared to [51]) of the samples in this work is related to the prevalence of titanium oxide inclusions in the material. Figure 21: SE-SEM image and EDS point analysis of Titanium oxide inclusion found in Test 2 fracture surface. 3.4. Summary The objective of this chapter was to conduct a thorough analysis of the bead deposition characteristics, microstructure, and mechanical properties of 18Ni-300 maraging steel printed using a micro-spot L-DED system. Based on the findings, the following conclusions can be drawn: 1. Powder feed rate significantly impacted bead AR and D due to the small 0.3 mm laser spot size (LSS). This spot size results in beads that are shorter and narrower compared to those 49 printed with larger LSSs, even with similar energy and mass densities, beads with higher average D. The average energy and mass densities to achieve ideal bead shapes (AR= 3 – 5 and D = 0.5 – 0.75) were found to be 74.5 J/mm2 and 0.018g/mm2, respectively. 2. Inspection of cuboid samples showed high relative density (>99%) but revealed shape and geometry defects. Externally, edge waviness, satellites, and unstable layer growth resulted from low powder capture/utilization efficiency associated with small LSSs. Internally, gas pores and irregular titanium oxide inclusions were caused by weld pool instability during L-DED and were also linked to the small LSS. Therefore, optimizing micro-spot L-DED systems should focus on balancing energy and mass input to the weld pool to avoid instability and improve powder capture efficiency, thereby reducing these defects. 3. The phase composition of the 18Ni-300 maraging steel was martensite with lesser amounts of retained austenite. A non-homogenous phase distribution and grain morphology were observed throughout the building height due to differing thermal histories and cooling rates. The bottom layers of cuboid samples, with higher cooling rates and re-heating, had more retained austenite and finer grains/cells. With lower cooling rates, the top layers of cuboids had less retained austenite and courser grains/cells. 4. The tensile properties of the as-built material were limited due to the presence of titanium oxide inclusions in the material, which were formed due to fluctuations in shielding gas flow and weld pool instabilities during printing—nevertheless, the tensile properties agreed with prior studies on L-DED of 18Ni-300 maraging steel. With suitable processing parameters outlined for printing high-density volumetric specimens using micro-spot L-DED, this system and the L-DED print settings can be applied to a practical repair application. The next chapter utilizes the L-DED settings according to parameter 50 Set 1 in Table 6 to perform a repair application. Although not all factors have been optimized for at this point (i.e., the elimination of titanium oxide inclusions which deteriorate tensile properties), this research commenced with applying these parameters to repair an HPDC insert. Discussion on how to eliminate these detrimental inclusions to improve the quality of the deposited material is provided in Chapter 5. 51 Chapter 4: Laser Directed Energy Deposition Repair of Maraging Steel High Pressure Die Cast Tooling Produced by Laser Powder Bed Fusion4 4.1. Introduction This chapter describes the design and execution of the L-DED repair process and the post-repair characterization and analysis. In this chapter, the suitable L-DED parameters identified in Chapter 3 for producing high-density cuboid samples are utilized to perform the repair. In the tool-making industry, AM technologies such as L-PBF have been beneficial for creating complex geometries, components with internal features (e.g., conformal cooling channels or intrinsic lubrication), and components from hard-to-machine alloys (e.g., M2, H11, and 18Ni-300) [6], [10], [11], [114]. Specifically, this technology has proven highly advantageous in creating dies and moulds for forging, injection moulding, and HPDC processes due to the ability to create unique designs that improve heat transfer conditions, prolonging tool life and reducing ecological harm (i.e., reduce CO2 emissions) [24], [75], [76]. However, even with these improvements, dies used in HPDC processes are still prone to wear and damage during service, resulting in the eventual failure and replacement [71], [72]. Heat checking due to thermal fatigue is one of the main failure modes for HPDC tools [115], in which thermally induced stresses produce microcracks near the working surface [116]. Once significant cracking is present, dies need to be replaced, incurring additional expense and revenue loss to the manufacturer through forced system downtime and part replacement costs 4 Part of the contents in this chapter have been adapted from: C. Paul, A. Jones, J. Vincic, N. Zaker, B. S. Amirkhiz, R. Mackay, G. Byczynski, A. Elsayed, L. Tobber, M. Benoit (2024), “Laser Directed Energy Deposition Repair of Maraging Steel High Pressure Die Cast Tooling Produced by Laser Powder Bed Fusion” (in preparation). 52 [117]. To combat high replacement costs, extend tool life, and reduce waste, traditional fusion welding techniques such as GTAW have been used to perform these types of repairs [71], [72], but this resulted in poor surface quality and required high heat input [118]. Fortunately, research has shown that repair by L-DED technology is beneficial as it allows for precise deposition of material on existing components with relatively minimal heat input and allows for deposition on non-planar surfaces [15], [24], [117]. This chapter aims to investigate the micro-spot L-DED repair of a high Ni maraging steel HPDC insert, originally produced by L-PBF, used for casting aluminum. The insert was supplied by Nemak, a Tier 1 automotive parts supplier [119]. This component failed due to heat check cracks forming on its surface during high-volume production. As the current study focuses on developing a repair process for a real part after service; the complex geometry and microstructure of the substrate create unique challenges that need to be overcome. Emphasis in the analysis of the repair is placed on the microstructural changes in the heat affected zone (HAZ) and the microstructural discrepancies between the deposited material in the fusion zone (FZ) and the base metal (BM) that is the original material of the HPDC insert. To the authors’ knowledge, this chapter outlines the first time one L-DED is used to repair a serviced part originally printed by L-PBF. However, a few examples exist of L-DED surface treating L-PBF parts in Errico et al. [120] and in the review by Asnafi [24]. 4.2. Methodology 4.2.1.Materials The HPDC insert that was repaired in this chapter had maximum height, width, and depth dimensions of 100.3 mm, 247.8 mm, and 176.2 mm, respectively, and a mass of approximately 17.1 kg. The insert was fabricated by L-PBF using an EOS printer and MS1 maraging steel 53 (nominal composition reported in Table 8), achieving a relative density of >99%. The insert was subjected to an aging heat treatment (HT) after printing to achieve an over-aged condition, with a targeted Rockwell C hardness (HRC) of 44 – 46 HRC. Given the industrial nature of this application, specifics related to the HPDC tooling geometry, L-PBF process parameters and scan strategy, and the aging HT cycle are proprietary. The HPDC insert under investigation in this chapter was placed into service and experienced ~5000 casting cycles before the heat-checking defect was identified, and the insert was removed for repair. Table 8: Nominal chemical compositions of powders used to print insert and for repair materials in wt%. Material Ni Co Mo Ti Cu Al Si Mn C Cr O N Fe Insert (MS1) 17.5 9.00 4.90 1.16 0.01 0.11 0.07 0.03 0.01 0.13 - - bal. Repair (18Ni-300) 18.4 9.07 4.86 0.82 0.1 0.07 0.03 0.02 0.02 0.01 0.01 0.01 bal. 4.2.2.Repair Process The overall process for repairing the cracked insert was conducted in six stages, as summarized in Figure 22. The basic steps and general details provided in the following paragraphs, describing the repair methodology, are adaptable and could be applied to various metals and their components. The specific details, however, such as the exact L-DED parameters (e.g., laser power, scan speed, powder feed rate, etc.), repair volume, and tool path planning, would all have to be tailored to the specific application. 54 Figure 22: Repair process methodology outlined over six stages. In Stage 1, the heat check crack was fractured to reveal the heat check crack surface (i) and the notch-sensitized fracture surface (ii). Nemak provided these images. Stage 2 required a multi-step machining process, starting with ball-milling (left image) and finishing with angled end-milling to create the landing B (right image) to avoid collisions during the repair process. Stage 1 – Defect identification: The location and extent of heat check cracking in the insert were identified. The prior independent analysis had already identified the location of the hairline crack (crack width < 1 mm), and fractography analysis revealed the maximum depth of the crack was 6 mm at the deepest point (according to region i in Figure 22-1). Identifying the crack depth was essential to remove the entire crack in Stage 2. Stage 2 – Defect removal: A manual vertical mill with a 6.35 mm (1/4”) ball mill cutter (Figure 22-2 left image) and 4.76 mm (3/16”) end mill cutter were used to progressively remove the crack, resulting in a groove in the HPDC insert that would be subsequently re-filled by L-DED in Stage 5 (Figure 22-2 right image). After creating the initial groove to remove the crack, the end mill was used to create 15° relief angles on the sides of the groove (labelled A, Figure 22-2 right image) as well as a 45° landing (labelled B, Figure 22-2 right image), which was made to avoid collisions between the tooling and laser deposition head during the repair process. The base shape of the groove was U-shaped to avoid any hard corners in the groove that could limit material fusion and bonding during L-DED [14]. 55 Stage 3 – Repair volume modelling: The repair volume to fill the groove created in Stage 2 was created and optimized using reverse engineering techniques. A Hexagon Absolute Arm 83 series laser scanning system was used to create a point cloud of the insert geometry with the machined groove so that a new geometry could be constructed using a negative space approach. As the die was scanned, it was carefully rotated onto two faces to ensure that all the possible geometrical information could be captured. Hexagon PC-DMIS computer-aided design (CAD) ++ metrology software was used to process the point cloud data and turn it into a mesh. The mesh was then compared with the original HPDC insert CAD model and a Boolean operation using Materialise Magics to create the groove repair volume. Stage 4 – L-DED Planning: Planning for the L-DED process required tool path design using the groove repair volume created in Stage 3 and parameter selection. The LAWS250 L-DED system was used to repair the HPDC insert. The L-DED process parameters used for the repair were previously determined to result in > 99.5% density on the LAWS 250 system and are provided in Table 9. The scan strategy during printing was a bi-directional zig-zag pattern that rotated 90° every two layers. The proprietary slicing software within the LAWS250 system was used to create the tool path for the repair volume. The hatch percentage between adjacent tracks and the layer thickness values presented in Table 9 determined the number of weld tracks per layer and number of layers, respectively, for the tool path. The PowderRangeE M300 high nickel maraging steel (referred to as 18Ni-300 maraging steel) from Carpenter Additive Inc. was used as the infill material, which was chosen because of its similar composition to the MS1 steel used to fabricate the HPDC insert (Table 8). 56 Table 9: L-DED printing parameters for the repair process. Parameter Value Laser Power 375 W Powder Feed rate 5 g/min Scan speed 750 mm/min Spot size 0.3 mm Hatch % 30% Layer Thickness 0.35 mm Argon Shielding Gas Flow 9 L/min Argon Carrier Gas flow 3 L/min Stage 5 – Part repair: Before the repair of the HPDC insert began, the insert was mounted to a custom positioning plate using bolts and bolt holes manufactured into the HPDC insert. The positioning plate was mounted in the LAWS250 build chamber, and the build surface for the repair was located using the in-line laser camera. This allowed the repair volume to be positioned to the highest level of accuracy possible in the LAWS250 slicing environment. With this achieved, the continuous L-DED repair process began (Figure 22-5) and was completed in under 7 minutes. The entire repair volume was 512.3 mm3. Stage 6 – Post-Repair Inspection: The repaired surface was visually inspected, and a laser scanning system was used to create a point cloud of geometrical data (similar to Stage 3) of the newly deposited material in the groove. This point cloud data was compared to the original CAD model of the HPDC insert to determine the geometrical accuracy of repair and any rise of part distortion due to the heat input. After external analysis, microstructural analysis samples were extracted from the repair region with wire electron discharge machining (EDM). 4.2.3.Microstructure analysis After wire EDM, the samples were prepared using typical metallographic techniques to achieve a mirror finish on the sample before etching in 5 vol% nital for two minutes. Optical microscopy (OM) was performed using a Keyence 970F to assess the characteristics of the weld pools in the fusion zone (FZ) (e.g., solidification structure, pores, and other defects). A Tescan 57 Mira Scanning Electron Microscope (SEM) with secondary electron (SE), backscatter electron (BSE), Oxford instruments X-Max energy dispersive x-ray spectroscopy (EDS) and Oxford Instruments Nordlys electron backscatter diffraction (EBSD) detectors were used to characterize further the BM, FZ, and heat affected zones (HAZs). The SE and EDS imaging and elemental analysis were performed using a 15 kV accelerating voltage, while EBSD mapping was performed with a 30 kV accelerating voltage and 0.1 μm step size. Chemical analysis of EDS data was performed using AztecLive software, while AztecCrystal analysis software was used for post-processing and analysis of EBSD data. 4.2.4.Mechanical Property Assessment Since the repair process was performed on the true part geometry, extraction of tensile specimens from the repair was not feasible. Hence, the mechanical properties of the repair region were assessed using indentation-based methods. Vickers micro-hardness indentation using a Wilson VH3100 motorized indentation system was used to map the microhardness across the entire repair region, including regions of the parent material on either side of the repair (n = 384 indentations total). Indentations were made using 500 gf (HV0.5) and a 12 s dwell time, and all indentations were spaced 0.5 mm apart in both horizontal and vertical directions, following the 2.5× indentation diagonal spacing recommended in ASTM E84 [121]. 4.3. Results and Discussion 4.3.1.External Assessment of Repair Figure 23 shows photographed images and point-cloud displacement data of the groove created to remove the heat-check crack (Figure 23a and c) and the new surface created by the L-DED repair process (Figure 23b and d). The dark blue region in Figure 23c indicates that the groove is below the original geometry of the insert’s original dimensions by more than 0.2 mm. 58 The deposited material in the repair groove follows the original contour of the insert’s design well, albeit with a rougher surface. The point cloud data in Figure 23d shows a region of overbuilt material (red region in Figure 23d) on the top and front of the filled groove, with the greatest deviation from the original surface of the insert CAD model is 0.34 mm. An over-filled repaired region was a goal for the repair process geometry so that the excess material could be machined back to produce a smooth, accurate surface. In a production environment, this could be achieved through the use of a hybrid additive-subtractive L-DED system [122], [123]. Conversely, undergrowth in the repair region (blue region in Figure 23d) was found to be 0.41 mm at the top of the landing that was machined for access of the deposition head due to unstable layer growth during repair [95]. The rest of the repair (i.e., the green region in Figure 23d that is between the red and blue regions) was within ±0.05 mm of the part’s original geometry. The overall dimensions of the insert did not significantly change due to heat input during the repair (i.e., distortion was negligible). Figure 23: Photograph of repair groove after machining a) and after printing b), accompanied with dimensional colormaps created from point cloud data captured via laser scanning c) and d), respectively (top-down view of repair region). The numbers for the colour map legend are in mm units. 59 4.3.2.Internal Assessment of Defects Figure 24 shows a low magnification backscatter electron (BSE) image of the repair region with several defects indicated. BSE increased the contrast of the image making it easier to see pores and defects in the chosen magnification. Unfortunately, several defects can be seen in the deposited 18Ni-300 (Figure 24), resulting in the relative density of the deposited material in this region being approximately 98.5%. A substantial lack of fusion between the deposited material and the side of the groove can be seen on the right of the groove (labelled A in Figure 24). This lack of fusion can be attributed to two related causes: (i) incorrect slicing and tool path planning and (ii) non-optimized groove geometry. First, the inaccurate tool path position stemmed from a misalignment of the HPDC insert in the LAWS250 build volume and the placement of the repair volume model in the slicing software. Without an integrated point cloud system (i.e., laser scanning), the manual alignment and transformation of coordinate systems to position the tool path on the insert resulted in a small position error. Upon further analysis of defect A (Figure 24), the maximum horizontal distance in the gap is ~0.167 mm, meaning there is potentially 0.167 mm in positioning error. The deposition pattern (e.g., zig-zag, honeycomb, or contour) can influence the formation of defects in the deposited region. For example, a zig-zag pattern has been found to produce deposited material with more voids than a contour pattern [66]. Additionally, suppose the repair volume is not divisible by the hatch spacing of single tracks. In that case, the single tracks can overlap in a pattern resulting in voids, another potential explanation for defect A in Figure 24. Unfortunately, a contour pattern was unavailable in the LAWS250 slicing software; thus, a less optimal zig-zag pattern needed to be used. Second, optimizing the groove volume and shape is crucial for a repair application to ensure fusion between the part and the deposited material [14], 60 [66]. U-shaped grooves without adequate relief angles have been shown to prevent a weld bead from correctly forming by limiting access of the filler metal to the laser spot, leading to lack-of-fusion voids [124]; relief angles as low as 5° have been shown to improve fusion with the side-wall substantially [14]. Thus, the 15° relief angle used in the current study should have provided adequate access for the deposition head and powder. However, further optimization of the groove design in conjunction with tool path planning is required for complete fusion throughout the repair. Relief angles were also required to ensure the weld head had access to the repair location without collision, hence the need for the 45° landing in the repair groove (labelled B in Stage 2 - Figure 22). Small- and medium-sized spherical pores, 0.03 mm < d< 0.15 mm (labelled B in Figure 24) were found throughout FZ. Irregular, high aspect ratio inclusions were also found (labelled C in Figure 24), these inclusions can be in various orientations within the repair volume. Additionally, non-spherical pores/voids were found at the interface of the FZ and base metal (BM) (labelled D in Figure 26). Lastly, large and small pores (labelled E in Figure 24) were prevalent near the very bottom of the groove (the black dashed arc in Figure 24), where the repair process began. It is worth noting that the BM was of high quality, with only small defects, presumably gas pores, found in the BM in the region around the repair, as highlighted by the black dashed box in Figure 24. 61 Figure 24: SEM-BSE image of the entire repair region, with several defects indicated, including (A) large lack-of fusion pore, (B) spherical pores, (C) irregularly shaped defects containing a dark material, (D) non-spherical pores/voids at the groove edge, and (E) pores that were prevalent at the bottom of the groove. The black dashed box highlights porosity in the BM, while the black dashed arc indicates the bottom of the repair geometry. The build direction (BD) is also indicated. Figure 25a-f presents optical micrographs of defects found in FZ. Figure 25a and b are irregularly shaped voids found at or near the repair groove boundary, indicative of lack of fusion (labelled D in Figure 24). An additional non-spherical and spherical inclusion can be seen in Figure 25a; these inclusion defects were commonly seen along the groove boundary (also indicated by D labels in Figure 24). Figure 25c and d capture spherical and non-spherical pores at the bottom of the weld track (labelled D in Figure 24), indicative of key-hole porosity due to high energy density [61], [125]. Figure 25e and f are taken from the center of the FZ and present gas pores and inclusion defects. While the gas porosity and lack of fusion defects are unsurprising in the deposited 18Ni-300, given these are common defects during L-DED, the inclusion defects containing a secondary phase requires further investigation. 62 Figure 25: Array of defects identified in the repair region, including a) and b) lack-of-fusion defects at the FZ-BM interface, c) and d) gas pores at the boundary of the weld pool, and e) and f) inclusions and small circular gas pores. The black dashed lines represent the interface between the BM and FZ. SEM and EDS analysis of a typical inclusion observed in the FZ (i.e., Figure 25f) can be seen in Figure 26. The EDS analysis (point analysis in Figure 26a and elemental map in c) shows that this inclusion mainly consists of titanium (Ti) and oxygen (O), with a small amount of aluminum (Al). The higher magnification image (Figure 26b) reveals a distinct interface between the inclusion, which can be seen to contain cracks and a dendritic structure, whereas the matrix of the deposited material can be seen to have a fine cellular structure in this region. The jagged nature and lack of dimples on the fracture surface of the large crack in the inclusion indicates that this 63 material is brittle compared to the surrounding 18Ni-300 maraging steel in the as-built condition. These characteristics of the titanium oxide were seen in all the remaining inclusions present throughout the FZ. Figure 26: SEM-SE images of the Ti/Al – O inclusion a) at 1kx magnification and b) at 5kx magnification, with c) corresponding EDS maps for Ti, Al, O, and Fe. Oxides of Ti and Al, such as the one shown in Figure 26 have previously been observed in as-built 18Ni-300 maraging steel processed by L-DED (as presented in Chapter 3) and L-PBF technology [49]. The proposed mechanism for the formation of these inclusions is fluctuation in the shielding environment and spatter from melt pools during printing, resulting in oxidized particles adhered to the surface part surface, which were then re-melted in the subsequent layer [84]. Spatter from unstable weld pools is associated with a high energy density during printing (e.g., 107 J/mm2 used) and performing L-DED with a small laser diameter (e.g., 0.3 mm), leading to poor powder capture efficiency [84], [47]. Fortunately, the LAWS 250 L-DED system has an integrated weld monitoring camera that records the entire repair process, allowing identification of the cause of the defects. Figure 27a shows the deposition head and repair groove immediately 64 before the L-DED repair process, with the as-machined groove side wall free of particles (highlighted in the green box). Figure 27b shows a still frame from the welding camera 43 seconds into the repair process, in which spatter and/or reflected particles can be seen moving away from the melt pool. Several particles can also be seen on the groove wall, in a region corresponding to where many inclusions were identified in Figure 24 (D labels). Spatter and shielding environment fluctuations, combined with a narrow repair groove design, contribute to the frequency of these inclusion defects in the FZ, as the narrow groove contains the spatter in a smaller area, as observed in the weld camera recording. Therefore, repair process parameters must be further optimized to reduce gas and lack-of-fusion pores, improve stability of shielding gas flows, and minimize spatter. Additionally, these brittle inclusions indicate that non-destructive examination methods such as X-ray computed tomography may be a necessary step in the repair process. Figure 27: a) laser deposition head and repair groove before repair and b) spatter and adhered particles on groove wall after 43 sec of the repair process. The green box outlines the side wall region, showing a) the clean wall before deposition and b) the wall with spatter particles after a few layers. 4.3.3.Microstructure Assessment of BM and FZ Figure 28a shows the etched micrograph of the as-built 18ni-300 maraging steel in the FZ, indicating different scanning directions. Figure 28b at higher magnification shows that fine equiaxed cells are found in the center of each track. In contrast, elongated columnar cells are found 65 closer to and orthogonal to the track boundaries. The heterogenous nature of this solidified structure is typical without HTs to homogenize it and remove the cellular structure [126]. Figure 28: Etched micrographs of the a) as-built L-DED 18Ni300 in the FZ (with scanning directions indicated by the black arrow and dot) with b) the solidification structure at the edge and center of the weld track. The build direction (BD) is indicated in both micrographs. Figure 29 shows an SEM-BSE image and corresponding EDS elemental maps for the as-built 18Ni-300 in the FZ and the over-aged MS1 in the BM. A heterogeneous distribution of Ti and molybdenum (Mo) exists within the FZ due to microsegregation that occurred during solidification [127]. Ti and Mo were rejected into the remaining phase, as evidenced by their increased concentration at cell boundaries in Figure 29a and b. Ti coincides with the dark particles in Figure 29a while, Mo aligns with the brighter regions. Microsegregation of Ti and Mo is commonly observed in the as-built microstructure of 18Ni maraging steels processed with L-DED [38], [126], [128]. No evidence of microsegregation can be observed in the BM due to the refined microstructure, or the aging HT removing it. 66 Figure 29: (a, c) BSE images and (b, d) corresponding EDS element maps for (a, b) the FZ and (c, d) the BM. Figure 30a shows the phase map for the as-built 18Ni-300 steel in the FZ. Figure 30a shows the body-centered cubic (bcc) structure (98.2%), assumed to be martensite, with trace amounts of face-centered cubic (fcc) austenite (1.13%); this austenite is retained austenite (RA). The small amount of RA within the inter-cellular regions of the bcc structure exists due to the rapid solidification and cooling, limiting the austenite to martensite transformation [74]. The orientation of RA will depend on the orientation of the dendrites and how they grew during solidification [53]. Comparing these results to Figure 29, microsegregation of Mo and Ti elements stabilized austenite in the intercellular regions, as Mo and Ti reduce the martensite transformation temperature (i.e., austenite stabilizing elements) [53], [126], [128]. 67 Figure 30: As-built L-DED in FZ a) EBSD phase, b) IPF-Y with IPF heat map, and c) grain boundary maps and HT L-PBF in the BM of the HPDC insert d) EBSD phase, e) IPF-Y with IPF heat map, and f) grain boundary maps. Figure 30b and c show the IPF-Y grain orientations and boundaries, respectively. The orientation of the grains shown in the IPF-Y map (Figure 30b) have a preferred orientation direction towards <001>, as seen from the IPF heat map, with respect to the build direction. This is common for as-built conditions of AM-processed steels [6]. Additionally, there are distinct groups of grains, such as the one indicated in Figure 30b, with some grains orientated in the <001> direction and other surrounding grains in the <111> direction. Figure 30c shows the different grain boundaries of the FZ microstructure: L-MAGBs range from 5° to 20° (black lines), PAGBs range from 20° to 50° (white lines) [53], and lastly, high-angle grain boundaries HAGBs >50° (red lines), which are indicative of martensite blocks [53]. The distribution of L-MAGBs, PAGBs, and HAGBs are 51.9%, 3.15%, and 44.9%, respectively. Cross-referencing the group of grains indicated in Figure 30 b and c shows that distinctly orientated grains are grouped because they are 68 within the same PAGB. Lastly, although hard to see from the figure, the RA grains within the PAGBs have a different orientation (<101>) than the martensite. During solidification, austenite will form first and then transform to martensite through shear deformation as it cools, but the remaining austenite will maintain the prior orientation [36], [129]. Figure 30 (d – f) the phase, IPF-Y, and grain boundary maps for the BM are shown. The over-aged 18Ni-300 maraging steel in the BM has more austenite present in the microstructure at 21.6% (fcc), with 76.4% martensite (bcc), as shown by Figure 30d. The austenite in the BM is reverted austenite (γ). In the IPF-Y map for the BM (Figure 30e), the grains are orientated in a way that is more typical (compared to the material in the FZ) of a martensitic structure, and it has a preferred <111> orientation with respect to the build direction. In Figure 30f, the L-MAGB, PAGB, and HAGB distributions are 55.2%, 10.4%, and 34.4%, respectively. The black dashed line in Figure 30e represents a packet of blocks (area inside white lines), which contains the martensitic laths. Figure 30f shows the same martensite packet outlined by the black dashed line within the PAGBs. This is typical for the microstructure formations of martensite in maraging steel [97]. From the analysis, the two materials in the FZ and BM exhibit vastly different microstructural characteristics. Most notable is the formation of a small amount of RA in the FZ but an increased amount of γ in the BM. The low amount of RA observed in the L-DED repair region is commonly observed for 18Ni-300 in the as-built condition when processed by L-DED and L-PBF, as both processes have fast cooling rates [109], [126]. Additionally, aging of this grade of maraging steel (typically achieved with 500°C) – in addition to precipitating Ni3Ti, Ni3Mo, and Fe2Mo intermetallic particles in the martensitic structure [126] – reverts bcc martensite into fcc austenite, forming γ [85], [98]. This is highlighted by comparing fcc amounts in Figure 30b and d. 69 Additionally, during aging HT, the microstructure will experience grain recrystallization and coarsening [98], which further explains the differences seen in grain orientation (preferential orientation <111> in BM versus <001> in FZ) and boundaries shown in Figure 30b and c and Figure 30e and f, respectively. 4.3.4.Microstructure Assessment of HAZ Figure 31a is an OM image that shows the HAZ transitioning from BM to the FZ. The HAZ is made up of two regions, labelled HAZ1 and HAZ2, respectively. HAZ1 is a bright region directly adjacent to the FZ boundary that experiences an austenitizing-like heat treatment (reaching temperatures of 1000 – 1300 °C), imparting a martensitic structure once cooling to room temperature [102], [130], [131]. HAZ2 is a darker colour with bright regions (white arrows) throughout that experiences an aging-like heat treatment (reaching temperatures of 550 – 650 °C), imparting precipitates in this region, giving it its dark colour [102], [130], [131]. The bright regions in HAZ2 are dispersions of γ, turning it into a predominant two-phase region [102]. EBSD phase mapping in Figure 32 helps confirm these findings. Figure 31b shows the distinct interface between HAZ1 and the FZ material. Typically, the deposited material grows as columnar cells that transition to finer equiaxed cells ~50 – 100 μm away from the interface. 70 Figure 31: The HAZ interface with different regions labelled a) with b) the hard transition between the HAZ1 and FZ. The HAZ width measurements were measured horizontally according to the “w” label. The white arrows indicate bright particles of reverted austenite. The build direction (BD) is indicated in both micrographs. The HAZ region undergoes a unique thermal history due to the multi-pass nature of the L-DED repair process. This creates a HAZ shape that follows the beads' deposition pattern at the edge of the repair groove. HAZ2 is wider than HAZ1, with a mean width along the height of the groove of 349.7 μm vs. a mean width of 81.9 μm, respectively, and the average total HAZ thickness was 432 μm. A summary of the HAZ widths is shown in Table 10. The HAZ width observed for the current L-DED repair is significantly smaller than the HAZ widths reported for welding technologies that can be used in repair, such as gas tungsten arc welding (GTAW) and plasma arc welding (PAW). Laser welding produces a narrower HAZ, up to 10 times as narrow as compared to GTAW and PAW processes – welding 18Ni-300 due to a lower heat input required [132], [133], [134]. Minimizing change to components BM (e.g., growth of γ) is beneficial for repair because it could reduce post-processing steps as well as increase the integrity of the repair in an as-deposited state [135], [136]. Optimization of L-DED process parameters (e.g., laser power, scan speed, and deposition strategy) to minimize heat input while still maintaining enough energy for melting could reduce the size of the HAZ. 71 Table 10: HAZ1 and HAZ2 width data summary. The error represents the standard deviation from 15 measurements in each HAZ. HAZ Average Width (μm) Max Width(μm) Min Width (μm) 1 82 (±42.7) 202.9 36.4 2 349 (±38.9) 410.9 256 Figure 32 shows the HAZ region's phase, IPF-Y, and grain boundary maps. In Figure 32a, the distribution of bcc and fcc phases is shown, with 75.2% bcc and 22.2% fcc, although a significantly higher amount of fcc (γ) is observed in the HAZ2 region (black arrow in Figure 32a). In Figure 32b, a typical martensitic microstructure for direct aging can again be seen in the HAZ2 and BM, while a less typical one is seen in the HAZ1 and FZ. The white arrows in Figure 32b are pointing to two instances where epitaxial growth can be seen. The white arrows indicate the regions in Figure 32a and c. In Figure 32c, these regions of epitaxial growth are contained within the respective PAGBs, as indicated by the white-lined grain boundaries. Epitaxial growth between the substrate and the deposited material is significant for bonding strength at the interface region [66], [137]. Figure 32c displays interesting behaviour of grain boundaries as well. For example, in HAZ1, the HAG boundaries are more prevalent than HAZ2 and the BM. This can be attributed to grain recrystallization and rapid cooling associated with the high temperature reached in this region from the heat input for repair [138]. 72 Figure 32: a) EBSD phase map, b) IPF-Y map and c) grain boundary map at the L-PBF to L-DED interface. 4.3.5.Mechanical Properties Figure 33 presents the Vickers hardness map of the repair region. The groove widens at the top due to the 15° relief angle added to each side during groove machining. This map does not include the L-DED material on the extra 45° landing (labelled B in Stage 2, Figure 22). The black and white lines in the figure approximate the FZ and HAZ boundaries, respectively. The L-DED 73 repair region generally had hardness values ranging from 300 – 350 HV0.5, and regions of exceptionally low hardness correspond to locations where the automated indentation was performed on a defect (Figure 33-i). The average hardness of the FZ and BM were 351.2 HV0.5 and 458.5 HV0.5, respectively (Table 11). The HPDC insert was specified to have a hardness of 44 – 46 HRC (430 – 460 HV) before service; hence, the BM still meets the manufacturer's requirements after L-DED repair of the heat-checking crack. The BM hardness also falls within the range of reported values for over-aged and peak-aged L-PBF MS1, which ranges from 400 HV to 630 HV [105], [107], [109], [139], [140], [141]. The FZ hardness falls within reported values for as-built L-DED, ranging in hardness from 320 to 396 HV, depending on the parameters that are used for processing [46], [53], [73], [74], [126], [142]. Figure 33: Vickers hardness data (HV0.5): map around the repair region with the black and white dashed lines enclosing the HAZ, i – indent performed on a defect during the indentation process, scale bar representing 50 μm. Table 11: Summary of Vickers hardness data. The error represents the standard deviation from 141 indents for the FZ and 243 indents from the BM, respectively. Region Vickers Hardness (HV0.5) Equivalent Rockwell (HRC) BM 459 (±11.6) 46 – 47 FZ 351 (±69.1) 35 – 36 Like hardness mapping for welds, hardness mapping of L-DED repaired parts can provide insight into the locations of softer materials, which are more prone to plastic deformation. Research on welded 18Ni maraging steels has shown the BM is typically the hardest in the weld region 74 [130], [143]. In some instances, though, depending on the heat input during the welding process, peak hardness has been found in HAZ2, which occurs due to an age-hardening effect [130]. This is not observed in the hardness map of the repair region. However, as described earlier, the aging-HT of this 18Ni maraging steel significantly changes the microstructure, leading to mechanical performance changes. Typically, aging-HT will positively affect the strength and hardness properties but negatively impact the ductility of 18Ni-300 compared to the as-built state [126], [128], [139], [140]. Prior results have shown that in tensile testing of L-DED repaired maraging steel, failure of these specimens in the softer FZ and HAZ due to the softer material experiencing more deformation [130], [132], [144]. Fortunately, hardness discrepancies between the FZ and the BM can be combated through proper solutionizing and/or aging HTs of the entire region, creating uniform hardness across both regions [102]. This is a widespread practice for welding techniques that can be used to improve L-DED repair methodologies further. Coalescence of material and microstructural properties post-L-DED repair is important for repaired parts to return to service. Fortunately, L-DED additionally offers process customization, such as remelting processes and inter-track/layer delays that can enable in-situ HT, improving mechanical properties [65], [73]. Exploration in in-situ HT and post-repair HT for L-DED repair is required to improve the outcomes of L-DED repair processes [8]. 4.4. Summary L-DED technology can repair damaged tools, reduce replacement costs, and extend tool life. This chapter aimed to develop a micro-spot L-DED repair process for a HPDC insert, originally fabricated by L-PBF and damaged from heat check cracking during its service. From the repair process and analysis of the repair region, the following conclusions can be made: 75 1. The micro-spot L-DED repair process developed for the HPDC insert produced an accurately shaped repair volume in a complexly shaped HPDC insert without impacting the overall geometry. However, some processing and metallurgical aspects of the repair process remain to be optimized. 2. Pervasive titanium oxide inclusions in the FZ appeared due to the remelting of spatter particles generated from unstable weld pools during micro-spot L-DED repair. Non-optimal tool path planning (accurate placement and deposition strategy) to match the shape of the repair groove is also attributed to the lack-of-fusion defect at the edge of the repair groove. 3. The as-built L-DED repair material differs microstructurally from the over-aged L-PBF BM. Most prevalent was the presence of retained austenite (RA) in the FZ but the presence of reverted austenite (γ) in the BM. The grain orientation with respect to the build direction was also different, <111> direction for BM but <001> direction for the FZ. The distinctive characteristics of the FZ and BM microstructure resulted in the BM hardness being 100 HV harder than the deposited material in the FZ. 4. The HAZ exhibited a unique microstructure compared to the FZ and BM. HAZ1 experienced an austenizing-like HT during the repair process, imparting increased HAG boundaries and RA in the region. HAZ2 experienced an aging-like HT during the repair process, increasing the region's reverted austenite (γ). A proof of concept design process has been designed and executed, proving the feasibility of micro-spot L-DED for industrial repair applications of damaged HPDC inserts made through L-PBF. Further optimization of repair parameters (i.e., L-DED parameters, groove shape, scan strategy, etc.) is required to eliminate many defects identified in the FZ. Post-processing HT, or 76 even intrinsic HT, steps were also not investigated, which could increase the material hardness of the FZ to match the BM hardness. Micro-spot L-DED parameters could be further optimized to increase hardness as it is known that L-DED parameters heavily impact as-built material characteristics [6], [19]. HT of repaired components has not been extensively reviewed in the literature, and the effects of HT on the entire HPDC insert again (after repair) are still unknown. Simulations of the repair process, using tools such as Simufact, would also be beneficial for understanding thermal histories that produce the microstructural changes (e.g., recrystallization of HAZ1 and reversion of austinite in HAZ2) observed in the HAZ. 77 Chapter 5: Extended Results and Discussion The work presented in Chapter 3 and Chapter 4 defined a process parameter window according to bead shape for a micro-spot L-DED system before repairing and analyzing a HPDC insert. This chapter further discusses the results of improving micro-spot L-DED processing for repair. First, a discussion on the weld pool stability of different parameter combinations, assessed via visual inspection of ST bead surfaces, is presented to provide insight into how bead morphology affects parameter selection. Then, a short set of experiments regarding the effect of gas flow parameters on deposition characteristics supported the discussion of changing gas flows to improve processing conditions during micro-spot L-DED. Lastly, the stages in the repair process that are problematic and in need of improvement are discussed. 5.1. L-DED Weld Pool Stability A stable weld pool is a weld bead pool that maintains consistent shape and size, steady temperature distribution, and minimal disturbances (i.e., spatter) throughout the welding process [127]. A key advantage of fusion-based AM technologies and automated welding processes is that they allow for live monitoring of process conditions. Advanced L-DED systems with in-situ monitoring systems, such as infrared probes or thermal cameras, can monitor weld pools as they are being built [145]. The L-DED system can use the information from these monitoring systems in a closed-loop feedback system to tune L-DED parameters in situ to achieve or maintain certain conditions (melt pool size, temperature profile, spatter control, layer growth, etc.) during printing [146], [147], [148]. If in-situ monitoring is not possible (e.g., on the LAWS250 system used in this thesis), information about weld pool behaviour, and thus parameters, can be gathered after printing by visual inspection of the printed material [149], [150]. Visual inspection of the weld track surface to check for bead continuity, surface condition/morphology, and the presence of particles/artifacts 78 can provide insight into the behaviour of the weld pool [45], [151], [152]. These factors are used to re-assess the ST samples created in Chapter 3. Figure 34 shows the energy density and mass density (according to Equation 1 and Equation 2, respectively) scatter plots according to the category of weld pool stability (Good, Satisfactory, or Poor) for the ST data sets presented in Table 5 (60 ST experiments total, 45 initial and 15 refined experiments). Weld beads were categorized as Good if the surface was smooth, had minimal spatter, and had a consistent straight weld bead edge along the length in the scan direction. Satisfactory weld beads had increased amounts of spatter compared to the Good category and acceptable weld bead edge along the bead length. The Poor weld beads had the most spatter on the surface and the least consistent weld bead edge (i.e., degree of waviness) along the length of the bead. The data is grouped according to STs that meet AR and D criteria (Figure 34a) and STs that do not (Figure 34b). Figure 35 shows Good, Satisfactory, and Poor weld bead surfaces, respectively. The ST surfaces were assessed based on visual inspection of the weld bead surfaces using a Keyence VHX 970F digital microscope. Figure 34: Weld pool stability categories for STs that meet shape criteria a) and STs that do not meet the shape criteria b). Weld pool stability assessed based on inspection of weld bead surface. 79 Figure 35: Image of weld beads in terms of weld pool stability a) “Good”, b) “Satisfactory,” and c) “Poor” that correspond to labels i, ii, and iii, respectively, in Figure 34 a) above. Scan direction (SD) indicated in a) applies to all images. The Good bead in Figure 35a has a smooth surface, is straight along the length of the bead, and has minimal particles on its surface. The Satisfactory and Poor ST weld bead surfaces (Figure 35b and c, respectively) are more oxidized (presence of purple-blue colour) [46], [153] and have more particles present. Figure 35c is deemed Poor compared to Figure 35b because more particles are present on its surface and around the bead. Particles on the surface are indicative of spatter ejected from the current weld pool (or subsequent-printed adjacent weld pools) during processing; excessive spatter is indicative of poor weld pool or process stability [84], [154]. The subjective nature of this assessment process should be acknowledged, as diverse levels of AM/welding knowledge and experiences could change the assessment outcome. Thus, the outcome is dependent on who is doing the assessment. An objective and quantitative assessment method should be developed to evaluate weld bead stability for AM/L-DED post-printing evaluation. One method that could be employed is laser or optical profilometry to quantify surface roughness. In the entire data set (Figure 34), only three Good weld beads regarding weld pool stability (making up 5% of the data set) were identified. Satisfactory and Poor settings regarding weld pool stability make up the remaining 35% (21 weld beads) and 60% (36 weld beads) of the data set, respectively. Only two Good weld beads meet the AR and D criteria. With these additional selection criteria, the window of process parameters with appropriate bead geometries and good stability now consists of two parameters rather than the original eighteen. The new set of 80 parameters that could be used has a laser power of 200 W, a scan speed of1000 mm/min, and a powder feed rate of 3 g/min (compared to the old parameters: 375 W, 700 mm/min, and 5 g/min). Figure 36 shows that these parameters produce a ST that meets the AR and D criteria previously used in section 3.3.1 Single Track Experiments (i.e., 3<AR<5 and 0.5 <D<0.75). It is important to note, from Figure 35a and Figure 36, that the weld beads with these new parameters are smaller than the previous best parameters (375 W, 700 mm/min, and 5 g/min), which is shown in Figure 35b). These parameters were identified as best in Chapter 3 while considering bead shape criteria, and they were subsequently used in Chapter 4 for repair. Using different parameters that produce a more stable weld pool for printing parts and performing the repair process could eliminate titanium oxide inclusions and key-hole pores in the material. However, changing parameters would also result in new hatch spacing and layer thickness parameters, which must be determined before printing volumetric samples. Additionally, with the smaller tracks, more beads would be needed to make any volume compared to larger beads. This would be beneficial for creating more accurate surfaces and filling the volume during the repair process. Figure 36: Micrograph of an ST cross-section of new potential optimum parameters if weld pool stability is considered. P=200 W, v = 1000 mm/min, and 𝑓𝑝 = 3 g/min. The AR and D for this parameter set is provided. 5.2. Assessing CG and SG Settings for Micro-spot L-DED The use of different powder carrier gas (CG) and shielding gas (SG) flow rates for micro-spot L-DED systems, with localized shielding, could help to combat the issue of spatter particles 81 and oxidation and achieve more stable weld pools, as evident from the work of Jardon et al. [45]. The most important would be to increase the CG flow and reduce the SG flow so that the ratio of CG/SG is >1. However, a high enough SG flow should be maintained, given the previous issues with shielding fluctuation. Increasing the CG/SG ratio has been reported to significantly increase the powder capture efficiency during printing and improve the weld pool stability (i.e., reducing spatter) [45]. Low powder utilization is a significant issue, with small LSSs contributing to excessive printing spatter [47]. Rather than using CG=3 L/min and SG=9 L/min, which is a common CG/SG ratio for larger laser spot systems (i.e., LSSs= 1 – 3 mm), a CG = 5 L/min and an SG=2 L/min should be used, according to the advice by Jardon et al. [45]. To evaluate the influence that the CG/SG ratio has on the print quality and bead shape, a small set of experiments were conducted at three different CG/SG ratios. The original gas settings for the results in Chapter 3 and Chapter 4 have a CG/SG = 0.3 (CG=3 L/min and SG =9 L/min). These values were selected based on recommendations by the manufacturer and what was found in the literature. Two new CG/SG ratios were selected: CG/SG = 1 (CG=6 L/min and SG =6 L/min) and CG/SG = 3 (CG=9L/min and SG =3 L/min). These values were inspired by the work done by Jardon et al. [45]. Five parameter sets shown in Table 12 were used to evaluate the different CG/SG ratios with the LAWS250 system. These parameters were selected based on the previous experimental findings of bead shape in Chapter 3. Sets 1 and 5 are outside the optimal shape window defined in Figure 9 while sets 2, 3, and 4 are inside the window. Set 2 corresponds to the new optimal parameters presented in Figure 36 and Set 3 corresponds to the parameters that were used for performing the repair of the HPDC insert. 82 Table 12: The five parameter sets used to experiment with different CG and SG flow settings. Set Power (W) Scan Speed (mm/min) Powder Feed Rate (g/min) 1 425 1400 3 2 200 1000 3 3 375 700 5 4 325 1000 5 5 275 600 8 To compare the impact of different gas settings on spatter formation, weld bead quality, and bead shape, images from the weld monitoring camera (integral to the LAWS250 system) and images captured of bead surfaces and cross sections on the Keyence VHX 970F digital microscope were used. Figure 37 shows images from the weld monitoring camera that were captured during printing. Only Sets 2, 3, and 5 are shown because they represent the three different powder feed rates used. Through the qualitative analysis of video recordings, the presence of spatter during ST printing compared across the three CG/SG ratios for Sets 2, 3, and 5 show that a CG/SG=3 (Figure 37) produces less spatter during printing compared to the CG/SG=0.3 and CG/SG=1 settings. Also worth noting is that across Sets 2, 3, and 5, the amount of excess powder on the build plate is visually lower at a CG/SG=3 compared to printing with a CG/SG=0.3. This could indicate that higher CG/SG gas settings are better for powder capture efficiency, as found previously by Jardon et al. [45], but quantitative measures would need to be implemented to confirm this. 83 Figure 37: Images taken from the weld monitoring camera in the LAWS250 during printing to assess spatter according to parameter sets 2, 3, and 5 at CG/SG=0.3, CG/SG=1, and CG/SG=3. Similar to the approach to assess weld pool stability in section 5.1, Figure 38 shows the weld bead surfaces for all parameter sets at the three CG/SG ratios. When comparing the presence of particles and consistency of bead shape across the CG/SG ratios, the bead shape improves, and the number of particles on the surface reduces at a CG/SG=3 compared to a CG/SG=0.3 in Figure 38. This is especially obvious when Set 3 is considered, but this trend is common across all sets. Furthermore, the particles on the surface of Set 2 – which was already identified as having “Good” stability with CG/SG=0.3 in section 5.1 – reduce even more at a CG/SG=3 (Figure 38-Set 2), further improving the stability of this parameter set. Lastly, the purple/blue colour on the surface of the bead made by Set 4 at a CG/SG=0.3 disappears at a CG/SG=3. 84 Figure 38: Weld bead surface images for the five parameter sets at CG/SG=0.3, CG/SG=1, and CG/SG=3. The scale bar at the bottom of Set 5, CG/SG=3, image represents a distance of 700 μm and applies to all images. Figure 39 shows the cross-sectional images for each parameter set across the three CG/SG ratios. Additionally, Figure 40 shows the corresponding bead dimensions and shape (aspect ratio and dilution), measured in ImageJ, for each parameter set across the CG/SG ratios. No significant shape or dimension changes were observed across the three CG/SG settings. Internally, though, an increase in porosity is observed at CG/SG=3, compared to CG/SG=1 in Figure 39. For Set 3, pores are present at all CG/SG ratios (at increasing amounts), while for Set 4 and 5, pores do not exist until the CG/SG= 3. Figure 39: Cross-sectional bead micrographs according to parameter set at CG/SG=0.3, CG/SG=1, and CG/SG=3. The scale bar at the bottom of all Set5, CG/SG=3, image represents a distance of 400 μm and applies to all images. 85 Figure 40: Bead a) height, b) width, c) depth, d) aspect ratio, and e) dilution at different CG/SG ratios for the five parameter sets. CG and SG flows have important roles in L-DED systems. CG carries the powder from the hopper to the melt pool, and SG protects the melt pool from oxidation [155], [156]. However, clear guidelines on gas flow rate values for L-DED systems, let alone micro-spot L-DED systems, have not been defined [155]. Generally, the guidelines that do exist are: “SG should be set high enough to prevent oxidation and formation of oxides” [157], and “CG should be set high so that a wide spread of powder is avoided at the focal distance but not too high to avoid excessive bouncing off of the substrate” [158]. Discussion on the effect of CG and SG flows on deposition characteristics is also conflicting. Jardon et al. [45] provides additional guidance on setting CG and SG values to achieve higher powder utilization efficiencies, but still, no clear ranges for CG or SG values are given. Some researchers say they do not affect bead formation [159] but others say they are critically important for bead surface morphology [160]. A review of the literature found that CG/SG ratios for L-DED of 18Ni-300 range from 0.1 to 1.8 (0.3 most common) with CG and SG values as high as 45 L/min and 28 L/min, respectively, for an underwater repair application, and as low as 3 L/min and 3 L/min, respectively [39], [40], [45], [73]. From these short experiments, two important observations can be made. First, increasing the CG/SG ratio reduced spatter during printing, which indicates better weld pool stability and 86 quality of material printed (i.e., minimizing chances of titanium oxide formation). Second, adjusting the CG/SG ratio has no major effect on the bead dimensions but does impact the presence of internal defects (e.g., gas porosity). With the application of micro-spot L-DED systems, these short experiments show that typical CG and SG settings are unsuitable for micro-spot L-DED systems, and new value ranges must be defined. The results of these short experiments are promising for future improvements for micro-spot L-DED as they agree with the findings of Jardon et al. [45], who found that the bead quality improves with increasing CG/SG ratio. 5.3. Areas to Improve HPDC Insert Repair Process The proof-of-concept repair process developed for the HPDC insert was outlined in Chapter 4 novel compared other repair studies in the literature (Table 3) as it used a real part with complex geometries, not plate geometries, which were manufactured with L-PBF and damaged with heat check cracking on its surface. The complex shape of the HPDC insert made defect removal, repair volume modelling, and tool path positioning – Stages 2, 3, and 4 in Figure 22 – difficult for the design of the repair process. This section discusses the challenges of the repair process and how they could be improved. The machining to remove the defect and create the repair groove was done on a manual 3-axis vertical mill. The size and shape of the repair groove in the insert have a significant role in the quality of the repair. The groove must be large enough to remove the defect but not too large to damage other aspects of the HPDC insert’s design. If designed properly, the repair groove could be optimized to minimize machining time, reduce material removal, and optimize shape for metallurgical bonding between the FZ and BM. Utilization of a multi-axis (>3) computer numerically controlled (CNC) machining center would make removing the defect and creating the repair groove easier and more optimizable. 87 The micro-spot L-DED repair process was also made difficult due to the capabilities of the LAWS250 system used. Firstly, the LAWS250 system does not have an integrated point cloud acquisition (PCA) system (i.e., laser scanning system) to capture important geometrical data. This would have made locating the working surface in the virtual slicing environment and positioning the repair volume much easier. This could reduce Stage 3 and Stage 4 of the repair methodology in Figure 22, to one single stage. Instead, with an external laser scanning system, geometrical and orientation data for repair volume had to be transferred manually to the LAWS250 slicing environment. Additionally, the positioning coordinates of the repair volume were calculated by hand, involving a known position on the positioning plate and a known position on the insert within the LAWS250 operating volume. An integrated PCA system would remove this manual positioning process and reduce the positioning error, which could eliminate the large lack of fusion defect (Figure 24). L-DED systems with integrated PCA systems are available and would greatly improve the workflows for repair processes [161]. Secondly, the LAWS250 system had limited scan strategies available in the slicing environment; thus, a zig-zag strategy needed to be used for the repair process. Unfortunately, though, this was identified as the inferior strategy compared to a contour strategy as outlined by the work done by Smail et al. [66]. The repair process would also be affected if different printing parameters were used (e.g., laser power, scan speed, and feed rate) according to considerations outlined in section 5.1 and 5.2 above. Given the results, the micro-spot L-DED repair process designed for the HPDC insert is a proof-of-concept process and was deemed successful. However, the development of this process was limited by the time and resources necessary to optimize the repair process to the highest degree (i.e., only one insert to attempt repair). 88 Chapter 6: Summary, Conclusion, and Future Work 6.1. Summary and Conclusion Fusion-based AM technologies (e.g., L-PBF and L-DED) are invaluable for the manufacturing industry in creating and repairing high-value components for HPDC operations and other manufacturing processes. Using fusion-based AM technology such as L-DED to repair high-value, complex casting inserts can increase tool life, reduce metal waste, and reduce operating costs. However, technical, economic, and environmental barriers still exist, slowing the adoption of this technology to repair tooling components. With advancements in AM technologies, such as the micro-spot L-DED method presented in this thesis, the possibilities and benefits of AM become greater but also more challenging to achieve. The advancement of technology requires the knowledge base to become broader and more specific to account for technological advancements. This thesis aimed to address technical barriers hindering the use of a novel micro-spot L-DED system for repairing damaged HPDC inserts, such as a suitable process parameter window and repair methodology. To achieve this goal, the research was divided into two segments that comprise this thesis's main research chapters. The first segment (Chapter 3) focused on the use of a novel micro-spot L-DED system to print 18Ni-300 maraging steel, understanding fundamental bead formation behaviour, microstructural characteristics, mechanical properties of printed material, and, most importantly, defining parameters for creating high-density material. The second segment (Chapter 4) focused on the development and execution of the repair process, utilizing suitable L-DED parameters identified in Chapter 3 and subsequent investigation into the microstructural and mechanical characteristics of the repair region, including the FZ, HAZ, and surrounding BM. Chapter 5 builds on the results and discussions of Chapters 3 and 4 by providing supplemental findings to support 89 claims made on how to improve micro-spot L-DED processing and outline future research directions for enhancing the repair process. From these chapters, the main conclusions of this thesis are: 1. Inspection of 18Ni-300 maraging steel single tracks (STs) printed with a micro-spot (LSS=0.3 mm) L-DED system showed inherently different bead formation characteristics than typical L-DED systems with larger LSSs. The 0.3 mm spot size results in shorter (bead height) and narrower (bead width) beads than those printed with larger diameter lasers, even with similar energy and mass densities. These characteristics influence the scan strategy planning used to design the repair volume tool path. The 0.3 mm spot size also creates beads with higher than typical dilution (D) values. Additionally, with a micro-spot L-DED system, the powder feed rate (fp) significantly impacts bead aspect ratio (AR) and D, having more influence than laser power (P) and scan speed (v) variables, which were previously reported as most significant on larger LSS systems. This is due to the low powder utilization associated with small LSSs. The ST shape characterization allowed for an operating window to be defined using average energy and mass densities to achieve the desired bead shape (AR = 3 – 5 and D = 0.5 – 0.75); these values were found to equal 74.5 J/mm2 and 0.018g/mm2, respectively. The parameters selected for repair were 375 W laser power, 700 mm/min scan speed, and a 5 g/min powder feed rate. 2. Inspection of cuboid samples showed high relative density (>99%). Externally, edge waviness, satellites, and unstable layer growth resulted from low powder capture/utilization efficiency associated with small LSSs. Internally, gas pores and irregular titanium oxide inclusions were caused by weld pool instability and fluctuations of the shielding environment during L-DED, also linked to the small LSS. The tensile properties of the as-90 built material were limited due to the presence of titanium oxide inclusions—nevertheless, the tensile properties agreed with prior studies on the L-DED processing of 18Ni-300 maraging steel. Microstructurally, the phase composition of the 18Ni-300 maraging steel was martensite with lesser amounts of retained austenite. The phase distribution and grain morphology observed along the build height of the sample were heterogeneous due to the different thermal histories and cooling rates at the bottom and top of cuboid samples. The sample bottom, with higher cooling rates and more re-heating, had more retained austenite and finer grains/cells. With lower cooling rates and no re-heating, the top layers had less austenite and courser grains/cells. 3. The micro-spot L-DED repair process developed for the HPDC insert produced an accurately shaped repair volume that followed the existing complex geometry of the HPDC insert surface without distortion of its overall geometry. In the FZ of the repair region, titanium oxide inclusions were prevalent due to the remelting of spatter particles that adhered to the surrounding nearby groove surfaces early on in the repair process. Non-optimal tool path planning (e.g., inaccurate placement and zig-zag deposition strategy) to match the shape of the repair groove also attributed to the large lack-of-fusion defect at the edge of the repair groove. The other side of the repair groove had a high degree of metallurgical bonding, as evident by the optical and scanning electron microscopy. 18Ni-maraging steel deposited in the FZ during the continuous repair had different microstructural characteristics compared to the over-aged MS1 maraging steel in the BM. The most prevalent was retained austenite (RA) in the FZ and reverted austenite (γ) in the BM. Additionally, the orientation of grains with respect to the build direction was also 91 different, with grains having a preferential <111> direction in the BM but grains having a preferential <001> direction in the FZ. 4. The HAZ is not an area that can be overlooked in L-DED-repaired components. Inspection of the HAZs revealed a unique microstructure compared to the FZ and BM. HAZ1 experienced an austenizing-like HT during repair, imparting increased HAG boundaries and RA in the region. HAZ2 experienced an aging-like HT during the repair, increasing the γ in this region. Vickers hardness (HV0.5) mapping of the FZ, HAZ, and BM showed that the FZ is softer than the BM. The different BM and FZ microstructure characteristics made the BM hardness ~100 HV0.5 harder than the deposited material in the FZ. 5. A greater understanding of weld bead stability, formation, and sensitivity to carrier gas (CG) and shielding gas (SG) settings was obtained through additional analysis of the weld bead surfaces (section 5.1) and performing short gas flow printing experiments (section 5.2). These were additional findings that were found through the analysis and explanation process of the results in Chapter 3. Inspection of weld bead surfaces provided insight into the stability of weld pools according to energy and mass densities, developing an understanding of parameters that result in titanium oxide inclusions. Considering stability criteria could change the scan strategy planning step of the repair process because the new ideal parameters (P=200 W, v = 1000 mm/min, and fp = 3 g/min) produced a much smaller bead than the original parameters. Gas flow experiments showed that the spatter during L-DED printing can be reduced with increasing CG/SG ratio. Additionally, increasing the CG/SG ratio has a negligible effect on bead dimensions but impacts the presence of internal defects. 92 6.2. Future Work This thesis demonstrated that micro-spot L-DED systems can repair heat-checking damaged HPDC inserts, but it also highlighted where certain parameters and aspects of the process need to be further explored for it to be fully optimized. The following are recommendations and comments for future work on micro-spot L-DED printing and repair: 1. For micro-spot L-DED, and in terms of addressing the lack of known optimal printing parameters, it was clear that only considering bead shape is insufficient to minimize defects in L-DED-produced 18Ni-300 maraging steels. Thus, suitable parameters also need to be defined in terms of stability (i.e., stable weld pools) to avoid excessive spatter and the formation of titanium oxide inclusions. With this multi-criteria optimization problem, a multi-criteria decision-making (MCDM) model, or something similar, could be developed and employed for ranking and selecting L-DED parameters. Aiming for more stable parameters would also improve the layer growth stability, avoiding undergrowth of parts and, thus, improving geometrical accuracy. Optimizing parameters to achieve a maximum hardness of 18Ni-300 maraging steel would benefit repair applications. 2. The effect of CG and SG parameters on deposition characteristics is not broadly defined or agreed on for L-DED technology. Initial experiments of ST bead printing at different CG/SG settings show that typical CG and SG values need to be updated for micro-spot systems. This warrants exploring the effect that varying CG and SG flow rates (at CG/SG≥1) has on weld bead dimensions, AR-D relationship, processing parameter windows, formation of defects in cubic volumes (particularly titanium oxide inclusions), and corresponding material properties (i.e., microstructure behaviour and mechanical properties). The focus of subsequent research should also be placed on minimizing 93 excessive D found in ST beads printed with a micro-spot system. Additionally, during the pursuit of this future work, an objective and quantitative measurement method should be employed to assess weld pool stability if no in-situ monitoring systems are available. 3. Optimization of groove geometry and scan strategy for the HPDC repair would improve the fusion between the FZ and BM, removing the large lack of fusion defect observed. This research did not perform groove optimization, as only one insert was available. Employing other slicing strategies, such as a contour slicing strategy, over the zig-zag scan strategy is also worth investing as it can remove the lack of fusion voids. Additionally, inter-layer pauses, no longer making the repair process continuous, could improve microstructure and hardness properties without needing post-repair HT. The vast customization available for micro-spot L-DED deposition strategies provides a large area for future research. 4. Finite element modelling of the L-DED process with Simufact was used in the pre-repair investigation (Chapter 3) to correlate thermal histories to microstructural behaviour according to location. Still, it was not used for the repair process due to processing power and time limitations. Simufact modelling could benefit the analysis of the HAZ as it would give insight into thermal histories experienced by the HPDC insert during the repair process. Similar to the pre-repair investigation, these thermal histories could be correlated to specific locations in the repair region, where unique microstructural phenomena and behaviour are observed. Simulations could also aid in defining process parameters for the repair process. 5. Employing a different L-DED system for repair, such as one with an integrated point cloud acquisition system, in-situ weld pool monitoring, and/or more than 3-axis movement, could improve the results. 94 6. Lastly, once the repair process is further optimized, an investigation will need to be done on how a repaired HPDC insert performs back in the production environment. 95 References [1]D. Herzog, V. Seyda, E. Wycisk, and C. Emmelmann, “Additive manufacturing of metals,” Acta Materialia, vol. 117, pp. 371–392, Sep. 2016, doi: 10.1016/j.actamat.2016.07.019. [2]N. Tuncer and A. Bose, “Solid-State Metal Additive Manufacturing: A Review,” JOM, vol. 72, no. 9, pp. 3090–3111, Sep. 2020, doi: 10.1007/s11837-020-04260-y. [3]T. DebRoy et al., “Additive manufacturing of metallic components – Process, structure and properties,” Progress in Materials Science, vol. 92, pp. 112–224, Mar. 2018, doi: 10.1016/j.pmatsci.2017.10.001. [4]Additive Manufacturing, Design, Functionally Graded Additive Manufacturing. ASTM International100 Barr Harbor Drive, PO Box C700, West Conshohocken, PA 19428-2959, 2021. doi: 10.1520/ISO/ASTMTR52912-EB. [5]C. Herron, M. Ivus, and A. Kotak, “Just Press ‘Print’: Canada’s Additive Manufacturing Ecosystem,” Information and Communications Technology Council (ICTC), Ottawa, ON, Canada, Mar. 2021. Accessed: Aug. 12, 2023. [Online]. Available: https://www.ictc-ctic.ca/wp-content/uploads/2021/04/Just-Press-Print-Canada%E2%80%99s-Additive-Manufacturing.pdf [6]P. Bajaj, A. Hariharan, A. Kini, P. Kürnsteiner, D. Raabe, and E. A. Jägle, “Steels in additive manufacturing: A review of their microstructure and properties,” Materials Science and Engineering A, vol. 772, no. November 2019, 2020, doi: 10.1016/j.msea.2019.138633. [7]J. Vinčić, A. Aversa, M. Lombardi, and D. Manfredi, “Processability and Microstructural Evolution of W360 Hot Work Tool Steel by Directed Energy Deposition,” Met. Mater. Int., vol. 30, no. 2, pp. 501–516, Feb. 2024, doi: 10.1007/s12540-023-01508-5. [8]C. Barr, R. A. R. Rashid, S. Palanisamy, J. Watts, and M. Brandt, “Examination of steel compatibility with additive manufacturing and repair via laser directed energy deposition,” Journal of Laser Applications, vol. 35, p. 11599, 2023, doi: 10.2351/7.0000952. [9]T. S. Tshephe, S. O. Akinwamide, E. Olevsky, and P. A. Olubambi, “Additive manufacturing of titanium-based alloys- A review of methods, properties, challenges, and prospects,” Heliyon, vol. 8, no. 3, p. e09041, Mar. 2022, doi: 10.1016/j.heliyon.2022.e09041. [10]J. Piekło and A. Garbacz-Klempka, “Use of Maraging Steel 1.2709 for Implementing Parts of Pressure Mold Devices with Conformal Cooling System,” Materials, vol. 13, no. 23, Art. no. 23, Jan. 2020, doi: 10.3390/ma13235533. [11]F. Bader, F. Hengsbach, K.-P. Hoyer, W. Homberg, and M. Schaper, “Intrinsically lubricated tool inserts for deep drawing applications generated by selective laser melting,” AIP Conference Proceedings, vol. 2113, no. 1, p. 170004, Jul. 2019, doi: 10.1063/1.5112720. [12]V. Laghi, M. Palermo, G. Gasparini, and T. Trombetti, “Computational design and manufacturing of a half-scaled 3D-printed stainless steel diagrid column,” Additive Manufacturing, vol. 36, no. August, p. 101505, 2020, doi: 10.1016/j.addma.2020.101505. [13]D. Jafari, T. H. J. Vaneker, and I. Gibson, “Wire and arc additive manufacturing: Opportunities and challenges to control the quality and accuracy of manufactured parts,” Materials & Design, vol. 202, p. 109471, Apr. 2021, doi: 10.1016/j.matdes.2021.109471. [14]B. Graf, A. Gumenyuk, and M. Rethmeier, “Laser Metal Deposition as Repair Technology for Stainless Steel and Titanium Alloys,” Physics Procedia, vol. 39, pp. 376–381, Jan. 2012, doi: 10.1016/j.phpro.2012.10.051. [15]K. Kanishka and B. Acherjee, “A systematic review of additive manufacturing-based remanufacturing techniques for component repair and restoration,” Journal of Manufacturing Processes, vol. 89, pp. 220–283, Mar. 2023, doi: 10.1016/j.jmapro.2023.01.034. 96 [16]G. Prashar, H. Vasudev, and D. Bhuddhi, “Additive manufacturing: expanding 3D printing horizon in industry 4.0,” International Journal on Interactive Design and Manufacturing, pp. 1–15, Jul. 2022, doi: 10.1007/s12008-022-00956-4. [17]C. van Sice and J. Faludi, “Comparing environmental impacts of metal additive manufacturing to conventional manufacturing,” Proceedings of the Design Society, vol. 1, no. August, pp. 671–680, 2021, doi: 10.1017/pds.2021.67. [18]H. Tupac-Yupanqui and A. Armani, “Additive Manufacturing of Functional Inconel 718 Parts from Recycled Materials,” Journal of Materials Engineering and Performance, vol. 30, no. 2, pp. 1177–1187, 2021, doi: 10.1007/s11665-020-05402-8. [19]H. Fayazfar et al., “A critical review of powder-based additive manufacturing of ferrous alloys: Process parameters, microstructure and mechanical properties,” Materials and Design, vol. 144, pp. 98–128, 2018. [20]M. Mani, K. W. Lyons, and S. K. Gupta, “Sustainability characterization for additive manufacturing,” Journal of Research of the National Institute of Standards and Technology, vol. 119, pp. 419–428, 2014, doi: 10.6028/jres.119.016. [21]A. Mecheter, F. Tarlochan, and M. Kucukvar, “A Review of Conventional versus Additive Manufacturing for Metals: Life-Cycle Environmental and Economic Analysis,” Sustainability, vol. 15, no. 16, Art. no. 16, Jan. 2023, doi: 10.3390/su151612299. [22]H. Hegab, N. Khanna, N. Monib, and A. Salem, “Design for sustainable additive manufacturing: A review,” Sustainable Materials and Technologies, vol. 35, p. e00576, 2023, doi: https://doi.org/10.1016/j.susmat.2023.e00576. [23]A. Kanyilmaz et al., “Role of metal 3D printing to increase quality and resource-efficiency in the construction sector,” Additive Manufacturing, vol. 50, no. November 2021, 2022. [24]N. Asnafi, “Tool and Die Making, Surface Treatment, and Repair by Laser-based Additive Processes,” Berg Huettenmaenn Monatsh, vol. 166, no. 5, pp. 225–236, May 2021, doi: 10.1007/s00501-021-01113-2. [25]P. Bidare, A. Jiménez, H. Hassanin, and K. Essa, “Porosity, cracks, and mechanical properties of additively manufactured tooling alloys: a review,” Advances in Manufacturing, vol. 10, no. 2, pp. 175–204, 2022, doi: 10.1007/s40436-021-00365-y. [26]W. M. Garrison and M. K. Banerjee, “Martensitic Non-Stainless Steels: High Strength and High Alloy,” Reference Module in Materials Science and Materials Engineering, 2018, doi: 10.1016/b978-0-12-803581-8.02519-4. [27]J. R. Davis, ASM Specialty Handbook: Tool Materials. ASM International, 1995. [28]A. P. Mourtiz, “Steels for aircraft structures,” in Introduction to Aerospace Materials, 2012, pp. 232–250. doi: 10.1533/9780857095152.232. [29]A. M. Hall and C. J. Slunder, “THE METALLURGY, BEHAVIOR, AND APPLICATION OF THE 18-PERCENT NICKEL MARAGING STEELS”. [30]T. H. G. Megson, “Chapter 11 - Materials,” in Aircraft Structures for Engineering Students (Sixth Edition), T. H. G. Megson, Ed., Butterworth-Heinemann, 2017, pp. 373–395. doi: 10.1016/B978-0-08-100914-7.00011-6. [31]D. A. Porter, K. E. Easterling, and M. Sherif, Phase Transformations in Metals and Alloys (Revised Reprint). Bosa Roca, UNITED STATES: Taylor & Francis Group, 2009. Accessed: Apr. 25, 2024. [Online]. Available: http://ebookcentral.proquest.com/lib/ubc/detail.action?docID=1449321 97 [32]H. Bhadeshia and R. Honeycombe, “Chapter 5 - Formation of Martensite,” in Steels: Microstructure and Properties (Fourth Edition), H. Bhadeshia and R. Honeycombe, Eds., Butterworth-Heinemann, 2017, pp. 135–177. doi: 10.1016/B978-0-08-100270-4.00005-6. [33]G. Thewlis, “Classification and quantification of microstructures in steels,” Materials Science and Technology, vol. 20, no. 2, pp. 143–160, Feb. 2004, doi: 10.1179/026708304225010325. [34]R. Tewari, S. Mazumder, I. S. Batra, G. K. Dey, and S. Banerjee, “Precipitation in 18 wt% Ni maraging steel of grade 350,” Acta Materialia, vol. 48, no. 5, pp. 1187–1200, Mar. 2000, doi: 10.1016/S1359-6454(99)00370-5. [35]V. K. Vasudevan, S. J. Kim, and C. M. Wayman, “Precipitation reactions and strengthening behavior in 18 Wt Pct nickel maraging steels,” Metallurgical transactions. A, Physical metallurgy and materials science, vol. 21 A, no. 10, pp. 2655–2668, 1990, doi: 10.1007/BF02646061. [36]G. I. Silman, “Compilative Fe – Ni phase diagram with author’s correction,” Met Sci Heat Treat, vol. 54, no. 3, pp. 105–112, Jul. 2012, doi: 10.1007/s11041-012-9463-x. [37]J. M. Pardal, S. S. M. Tavares, V. F. Terra, M. R. Da Silva, and D. R. Dos Santos, “Modeling of precipitation hardening during the aging and overaging of 18Ni–Co–Mo–Ti maraging 300 steel,” Journal of Alloys and Compounds, vol. 393, no. 1, pp. 109–113, May 2005, doi: 10.1016/j.jallcom.2004.09.049. [38]C. Félix-Martínez, J. Ibarra-Medina, D. A. Fernández-Benavides, L. A. Cáceres-Díaz, and J. M. Alvarado-Orozco, “Effect of the parametric optimization and heat-treatment on the 18Ni-300 maraging steel microstructural properties manufactured by directed energy deposition,” International Journal of Advanced Manufacturing Technology, vol. 115, no. 11–12, pp. 3999–4020, 2021, doi: 10.1007/s00170-021-07320-y. [39]G. Asala, O. T. Ola, and O. A. Ojo, “Effects of process variables on the quality and mechanical properties of 18%Ni-M350 maraging steel produced by direct energy deposition laser additive manufacturing,” Materials Science and Engineering: A, vol. 866, no. January, p. 144646, 2023, doi: 10.1016/j.msea.2023.144646. [40]M. W. Vaughan et al., “Development of a process optimization framework for fabricating fully dense advanced high strength steels using laser directed energy deposition,” Additive Manufacturing, vol. 67, p. 103489, Apr. 2023, doi: 10.1016/j.addma.2023.103489. [41]D.-G. Ahn, “Directed Energy Deposition (DED) Process: State of the Art,” Int. J. of Precis. Eng. and Manuf.-Green Tech., vol. 8, no. 2, pp. 703–742, Mar. 2021, doi: 10.1007/s40684-020-00302-7. [42]E. Toyserkani, A. Khajepour, and S. F. Corbin, Laser Cladding. Boca Raton: CRC Press, 2004. doi: 10.1201/9781420039177. [43]C. Felix-Martinez, L. D. Garcia-Salas, J. M. Gonzalez-Carmona, H. Ruiz-Luna, angel I. Garcia-Moreno, and J. M. Alvarado-Orozco, “Microstructure, hardness, and tribological performance of D2 tool steel fabricated by laser cladding using pulsed wave and substrate heating,” Optics and Laser Technology, vol. 175, 2024, [Online]. Available: http://dx.doi.org/10.1016/j.optlastec.2024.110862 [44]N. Rońda, K. Grzelak, M. Polański, and J. Dworecka-Wójcik, “The Influence of Layer Thickness on the Microstructure and Mechanical Properties of M300 Maraging Steel Additively Manufactured by LENS® Technology,” Materials, vol. 15, no. 2, 2022, doi: 10.3390/ma15020603. [45]Z. Jardon, J. Ertveldt, M. Hinderdael, and P. Guillaume, “Process parameter study for enhancement of directed energy deposition powder efficiency based on single-track geometry 98 evaluation,” Journal of Laser Applications, vol. 33, no. 4, p. 042023, 2021, doi: 10.2351/7.0000516. [46]S. Li, B. Chen, C. Tan, and X. Song, “Effects of oxygen content on microstructure and mechanical properties of 18Ni300 maraging steel manufactured by laser directed energy deposition,” Optics and Laser Technology, vol. 153, no. April 2021, p. 108281, 2022, doi: 10.1016/j.optlastec.2022.108281. [47]Y. Kong, L. Zhao, L. Zhu, and H. Huang, “The selection of laser beam diameter in directed energy deposition of austenitic stainless steel: A comprehensive assessment,” Additive Manufacturing, vol. 52, p. 102646, Apr. 2022, doi: 10.1016/j.addma.2022.102646. [48]P. Wołosz, A. Baran, and M. Polański, “The influence of laser engineered net shaping (LENSTM) technological parameters on the laser deposition efficiency and properties of H13 (AISI) steel,” Journal of Alloys and Compounds, vol. 823, p. 153840, May 2020, doi: 10.1016/j.jallcom.2020.153840. [49]B. Chen, Y. Huang, T. Gu, C. Tan, and J. Feng, “Investigation on the process and microstructure evolution during direct laser metal deposition of 18Ni300,” Rapid Prototyping Journal, vol. 24, no. 6, pp. 964–972, Jan. 2018, doi: 10.1108/RPJ-01-2018-0022. [50]K. Aleksandr, S. Ferdinando, R. Joel, C. Joel, M. Jordan, and J. Thomas, “Effect of direct energy deposition parameters on morphology, residual stresses, density, and microstructure of 1.2709 maraging steel,” International Journal of Advanced Manufacturing Technology, vol. 117, no. 3–4, pp. 1287–1301, 2021, doi: 10.1007/s00170-021-07635-w. [51]J. Gil, R. Seca, R. Amaral, O. Emadinia, A. De Jesus, and A. Reis, “Tensile Properties of As-Built 18Ni300 Maraging Steel Produced by DED,” Applied Sciences (Switzerland), vol. 12, no. 21, 2022, doi: 10.3390/app122110829. [52]L. Jiang, C. Chen, and M. Zhang, “Effect of Heat Treatment on the Properties of a High Alloying Maraging Steel Fabricated by Laser Metal Deposition,” Met. Mater. Int., vol. 29, no. 1, pp. 141–156, Jan. 2023, doi: 10.1007/s12540-022-01218-4. [53]E. A. Jägle, Z. Sheng, P. Kürnsteiner, S. Ocylok, A. Weisheit, and D. Raabe, “Comparison of Maraging Steel Micro- and Nanostructure Produced Conventionally and by Laser Additive Manufacturing,” Materials, vol. 10, no. 1, Art. no. 1, Jan. 2017, doi: 10.3390/ma10010008. [54]O. Emadinia, J. Gil, R. Amaral, C. Lopes, R. Rocha, and A. Reis, “Laser Deposited 18Ni300 Alloy Powder on 1045 Steel: Effect of Passes and Preheating on Microstructure,” Materials, vol. 15, no. 3, Art. no. 3, Jan. 2022, doi: 10.3390/ma15031209. [55]H. M. Zhu, J. W. Zhang, J. P. Hu, M. N. Ouyang, and C. J. Qiu, “Effects of aging time on the microstructure and mechanical properties of laser-cladded 18Ni300 maraging steel,” J Mater Sci, vol. 56, no. 14, pp. 8835–8847, May 2021, doi: 10.1007/s10853-021-05841-1. [56]S. Felicioni, A. Aversa, E. Librera, F. Bondioli, and P. Fino, “Directed energy deposition of 18NiM300 steel: effect of process and post processing conditions on microstructure and properties,” Science and Technology of Advanced Materials, vol. 25, no. 1, p. 2346071, Dec. 2024, doi: 10.1080/14686996.2024.2346071. [57]Y. Zhao et al., “Effect of beam energy density characteristics on microstructure and mechanical properties of Nickel-based alloys manufactured by laser directed energy deposition,” Journal of Materials Processing Technology, vol. 319, p. 118074, Oct. 2023, doi: 10.1016/j.jmatprotec.2023.118074. [58]H. Sohn, P. Liu, H. Yoon, K. Yi, L. Yang, and S. Kim, “Real-time porosity reduction during metal directed energy deposition using a pulse laser,” Journal of Materials Science & Technology, vol. 116, pp. 214–223, Jul. 2022, doi: 10.1016/j.jmst.2021.12.013. 99 [59]J. S. Kim, B. J. Kang, and S. W. Lee, “An experimental study on microstructural characteristics and mechanical properties of stainless-steel 316L parts using directed energy deposition (DED) process,” Journal of Mechanical Science and Technology, vol. 33, no. 12, pp. 5731–5737, 2019, doi: 10.1007/s12206-019-1116-1. [60]N. Shamsaei, A. Yadollahi, L. Bian, and S. M. Thompson, “An overview of Direct Laser Deposition for additive manufacturing; Part II: Mechanical behavior, process parameter optimization and control,” Additive Manufacturing, vol. 8, pp. 12–35, 2015, doi: 10.1016/j.addma.2015.07.002. [61]D. Svetlizky et al., “Directed energy deposition (DED) additive manufacturing: Physical characteristics, defects, challenges and applications,” Materials Today, vol. 49, pp. 271–295, Oct. 2021, doi: 10.1016/j.mattod.2021.03.020. [62]L. Guo, L. Zhang, J. Andersson, and O. Ojo, “Additive manufacturing of 18% nickel maraging steels: Defect, structure and mechanical properties: A review,” Journal of Materials Science & Technology, vol. 120, pp. 227–252, Sep. 2022, doi: 10.1016/j.jmst.2021.10.056. [63]E. M. Birger, G. V. Moskvitin, A. N. Polyakov, and V. E. Arkhipov, “Industrial laser cladding: current state and future,” Welding International, vol. 25, no. 3, pp. 234–243, Mar. 2011, doi: 10.1080/09507116.2010.540880. [64]R. A. Rahman Rashid et al., “Effect of clad orientation on the mechanical properties of laser-clad repaired ultra-high strength 300 M steel,” Surface and Coatings Technology, vol. 380, no. October, p. 125090, 2019, doi: 10.1016/j.surfcoat.2019.125090. [65]C. Barr et al., “Role of deposition strategy and fill depth on the tensile and fatigue performance of 300 M repaired through laser directed energy deposition,” International Journal of Fatigue, vol. 146, no. August 2020, p. 106135, 2021, doi: 10.1016/j.ijfatigue.2020.106135. [66]B. S. Smail et al., “Integrated approach to stainless steel 316L parts repair for pitting corrosion using laser metal deposition,” Journal of Manufacturing Processes, vol. 95, pp. 1–13, Jun. 2023, doi: 10.1016/j.jmapro.2023.04.007. [67]S. Jhavar, C. P. Paul, and N. K. Jain, “Micro-Plasma Transferred Arc Additive Manufacturing for Die and Mold Surface Remanufacturing,” JOM, vol. 68, no. 7, pp. 1801–1809, Jul. 2016, doi: 10.1007/s11837-016-1932-z. [68]O. Hentschel et al., “Experimental investigations of processing the high carbon cold-work tool steel 1.2358 by laser metal deposition for the additive manufacturing of cold forging tools,” Journal of Laser Applications, vol. 29, no. 2, 2017, [Online]. Available: http://dx.doi.org/10.2351/1.4983247 [69]R. Devine, C. Cullen, J. Foster, M. Kulakov, C. MacFadden, and S. Fitzpatrick, “Remanufacture of Hot Forging Dies By LMD-p Using a Cobalt Based Hard-Facing Alloy,” Berg Huettenmaenn Monatsh, vol. 166, no. 5, pp. 243–249, May 2021, doi: 10.1007/s00501-021-01108-z. [70]J. Grum and J. M. Slabe, “Possibility of introducing laser surfacing into maintenance service of die-casting dies,” Surface and Coatings Technology, vol. 180–181, pp. 596–602, Mar. 2004, doi: 10.1016/j.surfcoat.2003.10.101. [71]D. Klobčar, J. Tušek, B. Taljat, L. Kosec, and M. Pleterski, “Aging of maraging steel welds during aluminium alloy die casting,” Computational Materials Science, vol. 44, no. 2, pp. 515–522, Dec. 2008, doi: 10.1016/j.commatsci.2008.04.011. [72]D. Klobčar, L. Kosec, B. Kosec, and J. Tušek, “Thermo fatigue cracking of die casting dies,” Engineering Failure Analysis, vol. 20, pp. 43–53, Mar. 2012, doi: 10.1016/j.engfailanal.2011.10.005. 100 [73]S. Wang et al., “Investigation of on-site repair of 18Ni300 by underwater laser direct metal deposition technique,” Journal of Manufacturing Processes, vol. 80, no. June, pp. 909–919, 2022, doi: 10.1016/j.jmapro.2022.06.039. [74]F. Niu, H. Xing, F. Kan, G. Ma, X. Sun, and D. Wu, “Remanufacturing of 3D-Damaged Components Based on LDED Technology,” J. of Materi Eng and Perform, Jun. 2023, doi: 10.1007/s11665-023-08464-6. [75]M. Sode et al., “Tailoring flow behavior and heat transfer in tempering channels for high-pressure die casting—analysis of potentials of commercial static mixers and prospects of additive manufacturing,” Int J Adv Manuf Technol, vol. 125, no. 11, pp. 5463–5477, Apr. 2023, doi: 10.1007/s00170-023-10920-5. [76]A. Wiedenegger, J. Bruckwilder, and C. Deutsch, “Ecological and Economic Benefits of Additive Manufacturing in High Pressure Die Casting,” Berg Huettenmaenn Monatsh, vol. 166, no. 5, pp. 237–242, May 2021, doi: 10.1007/s00501-021-01110-5. [77]Z. C. Oter et al., “Benefits of laser beam based additive manufacturing in die production,” Optik, vol. 176, pp. 175–184, Jan. 2019, doi: 10.1016/j.ijleo.2018.09.079. [78]M. A. Kottman, “Additive Manufacturing of Maraging 250 Steels for the Rejuvenation and Repurposing of Die Casting Tooling,” Case Western Reserve University, 2015. Accessed: Jun. 12, 2024. [Online]. Available: https://etd.ohiolink.edu/acprod/odb_etd/etd/r/1501/10?clear=10&p10_accession_num=case1416854466 [79]J. Gil, R. Seca, R. Amaral, O. Emadinia, A. Reis, and A. de Jesus, “18Ni300 Maraging Steel Produced via Direct Energy Deposition on H13 Tool Steel and DIN CK45,” Key Engineering Materials, vol. 926, pp. 194–205, 2022, doi: 10.4028/p-8134oi. [80]D.-R. Eo, J. Yoon, and H. Lee, “Heterostructure effect at the interface of maraging steel deposited upon carbon steel via directed energy deposition,” Materials & Design, vol. 225, p. 111574, Jan. 2023, doi: 10.1016/j.matdes.2022.111574. [81]B. Zhu, L. Meng, Q. Hu, X. Zeng, X. Liu, and G. Xu, “Investigation of functionally graded materials for the repair of bainitic crossing noses using laser directed energy deposition: microstructure and performance,” Virtual and Physical Prototyping, vol. 18, no. 1, p. e2231905, Dec. 2023, doi: 10.1080/17452759.2023.2231905. [82]P. Sawczuk et al., “Regeneration of the Damaged Parts with the Use of Metal Additive Manufacturing—Case Study,” Materials, vol. 16, no. 10, Art. no. 10, Jan. 2023, doi: 10.3390/ma16103772. [83]“Optimising SEM Condition | EBSD Hints & Tips,” Oxford Instruments. Accessed: Aug. 15, 2024. [Online]. Available: https://www.ebsd.com/hints-and-tips/optimization-sem-condition [84]Y. Wang et al., “Tuning process parameters to optimize microstructure and mechanical properties of novel maraging steel fabricated by selective laser melting,” Materials Science and Engineering A, vol. 823, no. March, p. 141740, 2021, doi: 10.1016/j.msea.2021.141740. [85]S. A. R. Shamsdini et al., “A relationship between the build and texture orientation in tensile loading of the additively manufactured maraging steels,” Additive Manufacturing, vol. 41, no. November 2020, p. 101954, 2021, doi: 10.1016/j.addma.2021.101954. [86]Th. Schubert, W. Löser, S. Schinnerling, and I. Bächer, “Alternative phase formation in thin strip casting of stainless steels,” Materials Science and Technology, vol. 11, no. 2, pp. 181–185, Feb. 1995, doi: 10.1179/mst.1995.11.2.181. 101 [87]J. W. Fu, Y. S. Yang, J. J. Guo, and W. H. Tong, “Effect of cooling rate on solidification microstructures in AISI 304 stainless steel,” Materials Science and Technology, vol. 24, no. 8, pp. 941–944, Aug. 2008, doi: 10.1179/174328408X295962. [88]A. Mithal, N. Maharjan, and S. Idapalapati, “Enhanced cooling rates in laser directed energy deposition with interlayer peening,” Rapid Prototyping Journal, vol. 29, no. 6, pp. 1289–1298, Jan. 2023, doi: 10.1108/RPJ-11-2022-0395. [89]A. Aversa, G. Marchese, and E. Bassini, “Directed energy deposition of aisi 316l stainless steel powder: Effect of process parameters,” Metals, vol. 11, no. 6, 2021, doi: 10.3390/met11060932. [90]E. Toyserkani, A. Khajepour, and S. Corbin, “Laser Cladding Laser Cladding,” New York, vol. 11, no. 2, p. 221, Aug. 2017, doi: 10.1201/9781420039177. [91]S. Donadello, V. Furlan, A. G. Demir, and B. Previtali, “Interplay between powder catchment efficiency and layer height in self-stabilized laser metal deposition,” Optics and Lasers in Engineering, vol. 149, p. 106817, Feb. 2022, doi: 10.1016/j.optlaseng.2021.106817. [92]J. B. Ferguson, B. F. Schultz, A. D. Moghadam, and P. K. Rohatgi, “Semi-empirical model of deposit size and porosity in 420 stainless steel and 4140 steel using laser engineered net shaping,” Journal of Manufacturing Processes, vol. 19, pp. 163–170, Aug. 2015, doi: 10.1016/j.jmapro.2015.06.026. [93]F. M. Sciammarella and B. Salehi Najafabadi, “Processing Parameter DOE for 316L Using Directed Energy Deposition,” Journal of Manufacturing and Materials Processing, vol. 2, no. 3, Art. no. 3, Sep. 2018, doi: 10.3390/jmmp2030061. [94]A. Ascari, A. H. A. Lutey, E. Liverani, and A. Fortunato, “Laser Directed Energy Deposition of Bulk 316L Stainless Steel,” Lasers in Manufacturing and Materials Processing, vol. 7, no. 4, pp. 426–448, 2020, doi: 10.1007/s40516-020-00128-w. [95]M. Liu, A. Kumar, S. Bukkapatnam, and M. Kuttolamadom, “A Review of the Anomalies in Directed Energy Deposition (DED) Processes & Potential Solutions - Part Quality & Defects,” Procedia Manufacturing, vol. 53, pp. 507–518, Jan. 2021, doi: 10.1016/j.promfg.2021.06.093. [96]P. Kürnsteiner, M. B. Wilms, A. Weisheit, P. Barriobero-Vila, E. A. Jägle, and D. Raabe, “Massive nanoprecipitation in an Fe-19Ni-xAl maraging steel triggered by the intrinsic heat treatment during laser metal deposition,” Acta Materialia, vol. 129, pp. 52–60, May 2017, doi: 10.1016/j.actamat.2017.02.069. [97]S. Morito, X. Huang, T. Furuhara, T. Maki, and N. Hansen, “The morphology and crystallography of lath martensite in alloy steels,” Acta Materialia, vol. 54, no. 19, pp. 5323–5331, Nov. 2006, doi: 10.1016/j.actamat.2006.07.009. [98]L. Kučerová, K. Burdová, S. Jeníček, and I. Chena, “Effect of solution annealing and precipitation hardening at 250 • C-550 • C on microstructure and mechanical properties of additively manufactured 1.2709 maraging steel,” 2021, doi: 10.1016/j.msea.2021.141195. [99]Y. Z. Lek, C. Wang, X. Shen, Z. Chen, U. Ramamurty, and K. Zhou, “Additive manufacturing of corrosion-resistant maraging steel M789 by directed energy deposition,” Materials Science and Engineering A, vol. 857, no. August, p. 144032, 2022, doi: 10.1016/j.msea.2022.144032. [100]J. Zhang, J. Fan, J. Xu, D. Yang, Y. Peng, and K. Wang, “The effect of heat input on the microstructure and mechanical properties of 18Ni 300 maraging steel fabricated by arc directed energy deposition,” Materials Science and Engineering: A, vol. 884, p. 145545, Sep. 2023, doi: 10.1016/j.msea.2023.145545. [101]F. Deirmina, P. A. Davies, and R. Casati, “Effects of Powder Atomization Route and Post-Processing Thermal Treatments on the Mechanical Properties and Fatigue Resistance of 102 Additively Manufactured 18Ni300 Maraging Steel,” Advanced Engineering Materials, vol. 24, no. 4, p. 2101011, 2022, doi: 10.1002/adem.202101011. [102]B. Jose, M. Manoharan, A. Natarajan, N. R. Muktinutalapati, G. Madhusudhan Reddy, and S. D. Meshram, “Current research and developments in welding of 18% nickel maraging steel,” Proceedings of the Institution of Mechanical Engineers, Part L: Journal of Materials: Design and Applications, vol. 237, no. 6, pp. 1295–1318, Jun. 2023, doi: 10.1177/14644207221142516. [103]C. Liu, Z. Zhao, D. O. Northwood, and Y. Liu, “A new empirical formula for the calculation of MS temperatures in pure iron and super-low carbon alloy steels,” Journal of Materials Processing Technology, vol. 113, no. 1, pp. 556–562, Jun. 2001, doi: 10.1016/S0924-0136(01)00625-2. [104]Y. Yao et al., “Microstructural heterogeneity and mechanical anisotropy of 18Ni-330 maraging steel fabricated by selective laser melting: The effect of build orientation and height,” Journal of Materials Research, vol. 35, no. 15, pp. 2065–2076, Aug. 2020, doi: 10.1557/jmr.2020.126. [105]K. Chadha, Y. Tian, P. Bocher, J. G. Spray, and C. Aranas, “Microstructure Evolution, Mechanical Properties and Deformation Behavior of an Additively Manufactured Maraging Steel,” Materials, vol. 13, no. 10, Art. no. 10, Jan. 2020, doi: 10.3390/ma13102380. [106]G. Krauss, “Martensite in steel: strength and structure,” Materials Science and Engineering: A, vol. 273–275, pp. 40–57, Dec. 1999, doi: 10.1016/S0921-5093(99)00288-9. [107]M. Król, P. Snopiński, J. Hajnyš, M. Pagáč, and D. Łukowiec, “Selective laser melting of 18Ni-300 maraging steel,” Materials, vol. 13, no. 19, 2020, doi: 10.3390/MA13194268. [108]S. Amirabdollahian et al., “Towards controlling intrinsic heat treatment of maraging steel during laser directed energy deposition,” Scripta Materialia, vol. 201, p. 113973, Aug. 2021, doi: 10.1016/j.scriptamat.2021.113973. [109]F. Zhou, R. Wu, W. Xie, and L. Zhang, “Effect of aging treatment on microstructure and properties of additively manufactured maraging steel,” Ironmaking & Steelmaking, Oct. 2020, Accessed: Jan. 19, 2024. [Online]. Available: https://www.tandfonline.com/doi/full/10.1080/03019233.2019.1651022 [110]P. Kürnsteiner, M. B. Wilms, A. Weisheit, P. Barriobero-Vila, E. A. Jägle, and D. Raabe, “Massive nanoprecipitation in an Fe-19Ni-xAl maraging steel triggered by the intrinsic heat treatment during laser metal deposition,” Acta Materialia, vol. 129, pp. 52–60, May 2017, doi: 10.1016/j.actamat.2017.02.069. [111]J. Song et al., “Effect of heat treatment on microstructure and mechanical behaviours of 18Ni-300 maraging steel manufactured by selective laser melting,” Optics & Laser Technology, vol. 120, p. 105725, Dec. 2019, doi: 10.1016/j.optlastec.2019.105725. [112]S. Yin et al., “The influence of aging temperature and aging time on the mechanical and tribological properties of selective laser melted maraging 18Ni-300 steel,” Additive Manufacturing, vol. 22, pp. 592–600, Aug. 2018, doi: 10.1016/j.addma.2018.06.005. [113]H. Bhadeshia and R. Honeycombe, “Chapter 11 - The Embrittlement and Fracture of Steels,” in Steels: Microstructure and Properties (Fourth Edition), H. Bhadeshia and R. Honeycombe, Eds., Butterworth-Heinemann, 2017, pp. 303–341. doi: 10.1016/B978-0-08-100270-4.00011-1. [114]J. Sjöström, U. Fager, and G. Broberg, “Impact of New Design in a Plastic Injection Mold Made by Additive Manufacturing,” Berg Huettenmaenn Monatsh, vol. 167, no. 9, pp. 422–427, Sep. 2022, doi: 10.1007/s00501-022-01265-9. [115]J. Sjöström and J. Bergström, “Thermal fatigue in hot-working tools,” Scandinavian Journal of Metallurgy, vol. 34, no. 4, pp. 221–231, 2005, doi: 10.1111/j.1600-0692.2005.00719.x. 103 [116]P. Solgi, M. Chenarani, A. R. Eivani, M. Ghosh, V. Kumar, and H. R. Jafarian, “Heat checking as a failure mechanism of dies exposed to thermal cycles: A review,” Journal of Materials Research and Technology, vol. 26, pp. 865–895, Sep. 2023, doi: 10.1016/j.jmrt.2023.07.170. [117]S. Jhavar, C. P. Paul, and N. K. Jain, “Causes of failure and repairing options for dies and molds: A review,” Engineering Failure Analysis, vol. 34, pp. 519–535, Dec. 2013, doi: 10.1016/j.engfailanal.2013.09.006. [118]A. Saboori, A. Aversa, G. Marchese, S. Biamino, M. Lombardi, and P. Fino, “Application of Directed Energy Deposition-Based Additive Manufacturing in Repair,” Applied Sciences, vol. 9, no. 16, Art. no. 16, Jan. 2019, doi: 10.3390/app9163316. [119]“Nemak.” Accessed: Jul. 01, 2024. [Online]. Available: https://www.nemak.com/about-us/?sc=0 [120]V. Errico, P. Posa, M. Mazzarisi, A. Angelastro, and S. L. Campanelli, “Effects of Laser-Deposited Maraging Steel on L-PBF 316L Component,” Metals, vol. 12, p. 1669, Oct. 2022, doi: 10.3390/met12101669. [121]E04 Committee, Test Method for Microindentation Hardness of Materials. doi: 10.1520/E0384-17. [122]R. E. Petruse and M.-C. Langa, “Enhancing Metal Forging Tools and Moulds: Advanced Repairs and Optimisation Using Directed Energy Deposition Hybrid Manufacturing,” Applied Sciences, vol. 14, no. 2, Art. no. 2, Jan. 2024, doi: 10.3390/app14020567. [123]G. A. Barragan De Los Rios, R. Ferreira, F. E. Mariani, E. J. da Silva, and R. T. Coelho, “Study of the surface roughness of a remanufactured bimetallic AISI 1045 and 316L SS part obtained by hybrid manufacturing (DED/HSM),” Int J Adv Manuf Technol, vol. 124, no. 9, pp. 3185–3199, Feb. 2023, doi: 10.1007/s00170-022-09179-z. [124]T. Yang, J. Liu, Y. Zhuang, K. Sun, and W. Chen, “Studies on the formation mechanism of incomplete fusion defects in ultra-narrow gap laser wire filling welding,” Optics & Laser Technology, vol. 129, p. 106275, Sep. 2020, doi: 10.1016/j.optlastec.2020.106275. [125]F. Lu, X. Li, Z. Li, X. Tang, and H. Cui, “Formation and influence mechanism of keyhole-induced porosity in deep-penetration laser welding based on 3D transient modeling,” International Journal of Heat and Mass Transfer, vol. 90, pp. 1143–1152, Nov. 2015, doi: 10.1016/j.ijheatmasstransfer.2015.07.041. [126]L. Jiang, C. Chen, and M. Zhang, “Effect of Heat Treatment on the Properties of a High Alloying Maraging Steel Fabricated by Laser Metal Deposition,” Met. Mater. Int., vol. 29, no. 1, pp. 141–156, Jan. 2023, doi: 10.1007/s12540-022-01218-4. [127]S. Kou, Welding Metallurgy (3rd Edition). John Wiley & Sons, 2021. [Online]. Available: https://app.knovel.com/hotlink/toc/id:kpWME00006/welding-metallurgy-3rd/welding-metallurgy-3rd [128]J. Jeong et al., “Mechanical properties of lamellar-structured 18Ni300 maraging steel manufactured via directed energy deposition,” Materials Science and Engineering: A, vol. 892, p. 146031, Feb. 2024, doi: 10.1016/j.msea.2023.146031. [129]J. Zhang, J. Fan, J. Xu, D. Yang, Y. Peng, and K. Wang, “The effect of heat input on the microstructure and mechanical properties of 18Ni 300 maraging steel fabricated by arc directed energy deposition,” Materials Science and Engineering: A, vol. 884, p. 145545, Sep. 2023, doi: 10.1016/j.msea.2023.145545. [130]L. Fanton, A. J. Abdalla, and M. S. Fernades de Lima, “Welding Journal | September 2014”. 104 [131]“Fracture toughness and microstructure of the heat-affected zone of a welded maraging steel.” Accessed: Jun. 05, 2024. [Online]. Available: https://www.tandfonline.com/doi/epdf/10.1179/030716974803288121?needAccess=true [132]L. Subashini, K. V. P. Prabhakar, S. Ghosh, and G. Padmanabham, “Comparison of laser-MIG hybrid and autogenous laser welding of M250 maraging steel thick sections—understanding the role of filler wire addition,” Int J Adv Manuf Technol, vol. 107, no. 3, pp. 1581–1594, Mar. 2020, doi: 10.1007/s00170-020-05113-3. [133]C. Magadum, R. K. Gupta, P. Senthil, M. Manohar, and S. Biju, “Study on Effect of Gouge Depth on the Microstructure and Tensile Properties of 18%Ni Maraging Steel Welds Made Through Hot Wire GTA Welding,” J. Inst. Eng. India Ser. D, vol. 104, no. 2, pp. 785–794, Dec. 2023, doi: 10.1007/s40033-022-00430-1. [134]P. R. Sakai et al., “Comparison of Mechanical and Microstructural Characteristics in Maraging 300 Steel Welded by three Different Processes: LASER, PLASMA and TIG,” Procedia Engineering, vol. 114, pp. 291–297, Jan. 2015, doi: 10.1016/j.proeng.2015.08.071. [135]L. Subashini, K. V. P. Prabhakar, R. C. Gundakaram, S. Ghosh, and G. Padmanabham, “Single Pass Laser-Arc Hybrid Welding of Maraging Steel Thick Sections,” Materials and Manufacturing Processes, vol. 31, no. 16, pp. 2186–2198, Dec. 2016, doi: 10.1080/10426914.2016.1221099. [136]B. Jose, M. Manikandan, N. Arivazhagan, N. R. Muktinutalapati, and G. Madhusudhan Reddy, “Development of a Novel Welding Technique With Reduced Heat-Input by Employing Double-Pulsed Gas Metal Arc Welding for Aerospace Grade 18% Ni Maraging Steel,” Journal of Manufacturing Science and Engineering, vol. 145, no. 021002, Sep. 2022, doi: 10.1115/1.4055248. [137]V. Braga et al., “Microstructural and mechanical aspects of laser metal deposited H13 powder for die repair,” Materials Today Communications, vol. 29, p. 102945, Dec. 2021, doi: 10.1016/j.mtcomm.2021.102945. [138]Z.-F. Hu, D.-F. Mo, C.-X. Wang, G.-Q. He, and C.-S. Chen, “Different Behavior in Electron Beam Welding of 18 Ni Co-free Maraging Steels,” J. of Materi Eng and Perform, vol. 17, no. 5, pp. 767–771, Oct. 2008, doi: 10.1007/s11665-007-9190-4. [139]G. Huang, K. Wei, J. Deng, and X. Zeng, “High power laser powder bed fusion of 18Ni300 maraging steel: Processing optimization, microstructure and mechanical properties,” Materials Science and Engineering: A, vol. 856, p. 143983, Oct. 2022, doi: 10.1016/J.MSEA.2022.143983. [140]R. Casati, J. N. Lemke, A. Tuissi, and M. Vedani, “Aging Behaviour and Mechanical Performance of 18-Ni 300 Steel Processed by Selective Laser Melting,” Metals, vol. 6, no. 9, Art. no. 9, Sep. 2016, doi: 10.3390/met6090218. [141]Z. Zhao et al., “Texture dependence on the mechanical properties of 18Ni300 maraging steel fabricated by laser powder bed fusion,” Materials Characterization, vol. 189, p. 111938, Jul. 2022, doi: 10.1016/J.MATCHAR.2022.111938. [142]J. Gil, R. Seca, R. Amaral, O. Emadinia, A. De Jesus, and A. Reis, “Tensile Properties of As-Built 18Ni300 Maraging Steel Produced by DED,” Applied Sciences, vol. 12, no. 21, Art. no. 21, Jan. 2022, doi: 10.3390/app122110829. [143]J. R. Joshi, M. Potta, K. Adepu, R. K. Katta, and M. R. Gankidi, “A comparative evaluation of microstructural and mechanical behavior of fiber laser beam and tungsten inert gas dissimilar ultra high strength steel welds,” Defence Technology, vol. 12, no. 6, pp. 464–472, Dec. 2016, doi: 10.1016/j.dt.2016.08.003. 105 [144]T. V. Fieger, M. F. Sattler, and G. Witt, “Developing laser beam welding parameters for the assembly of steel SLM parts for the automotive industry,” Rapid Prototyping Journal, vol. 24, no. 8, pp. 1288–1295, Jan. 2018, doi: 10.1108/RPJ-12-2016-0204. [145]Z. Tang et al., “A review on in situ monitoring technology for directed energy deposition of metals,” Int J Adv Manuf Technol, vol. 108, no. 11, pp. 3437–3463, Jun. 2020, doi: 10.1007/s00170-020-05569-3. [146]C. Kledwig, H. Perfahl, M. Reisacher, F. Brückner, J. Bliedtner, and C. Leyens, “Analysis of Melt Pool Characteristics and Process Parameters Using a Coaxial Monitoring System during Directed Energy Deposition in Additive Manufacturing,” Materials, vol. 12, no. 2, Art. no. 2, Jan. 2019, doi: 10.3390/ma12020308. [147]J. Mi et al., “In-situ monitoring laser based directed energy deposition process with deep convolutional neural network,” J Intell Manuf, vol. 34, no. 2, pp. 683–693, Feb. 2023, doi: 10.1007/s10845-021-01820-0. [148]S. Donadello, M. Motta, A. G. Demir, and B. Previtali, “Monitoring of laser metal deposition height by means of coaxial laser triangulation,” Optics and Lasers in Engineering, vol. 112, pp. 136–144, Jan. 2019, doi: 10.1016/j.optlaseng.2018.09.012. [149]G. Horváth, A. Körmöczi, T. Szörényi, and Z. Geretovszky, “Bead morphology of laser welded steel sheets and its correlation to the joint’s electrical and mechanical properties,” Journal of Laser Applications, vol. 35, no. 2, p. 022020, Apr. 2023, doi: 10.2351/7.0000996. [150]S. F. Nabavi, A. Farshidianfar, and H. Dalir, “A comprehensive review on recent laser beam welding process: geometrical, metallurgical, and mechanical characteristic modeling,” Int J Adv Manuf Technol, vol. 129, no. 11, pp. 4781–4828, Dec. 2023, doi: 10.1007/s00170-023-12536-1. [151]N. L. Do Vale, C. A. Fernandes, R. de A. Santos, T. F. A. Santos, and S. L. Urtiga Filho, “Effect of Laser Parameters on the Characteristics of a Laser Clad AISI 431 Stainless Steel Coating on Carbon Steel Substrate,” JOM, vol. 73, no. 10, pp. 2868–2877, Oct. 2021, doi: 10.1007/s11837-021-04835-3. [152]E. W. A. Figueredo et al., “Influence of Laser Beam Power and Scanning Speed on the Macrostructural Characteristics of AISI 316L and AISI 431 Stainless Steel Depositions Produced by Laser Cladding Process,” J. of Materi Eng and Perform, vol. 30, no. 5, pp. 3298–3312, May 2021, doi: 10.1007/s11665-021-05676-6. [153]C. Barr, R. a. R. Rashid, S. Palanisamy, N. Matthews, and M. Brandt, “Effect of oxygen shielding on the tensile and fatigue performance of 300M repaired through laser-directed energy deposition,” The Aeronautical Journal, vol. 127, no. 1318, pp. 2118–2124, Dec. 2023, doi: 10.1017/aer.2023.102. [154]T. Furumoto, K. Oishi, S. Abe, K. Tsubouchi, M. Yamaguchi, and A. T. Clare, “Evaluating the thermal characteristics of laser powder bed fusion,” Journal of Materials Processing Technology, vol. 299, no. September 2021, 2022, doi: 10.1016/j.jmatprotec.2021.117384. [155]A. Singh, S. Kapil, and M. Das, “A comprehensive review of the methods and mechanisms for powder feedstock handling in directed energy deposition,” Additive Manufacturing, vol. 35, p. 101388, Oct. 2020, doi: 10.1016/j.addma.2020.101388. [156]U. de Oliveira, V. Ocelík, and J. Th. M. De Hosson, “Analysis of coaxial laser cladding processing conditions,” Surface and Coatings Technology, vol. 197, no. 2, pp. 127–136, Jul. 2005, doi: 10.1016/j.surfcoat.2004.06.029. [157]D.-R. Eo, S.-H. Park, and J.-W. Cho, “Controlling inclusion evolution behavior by adjusting flow rate of shielding gas during direct energy deposition of AISI 316 L,” Additive Manufacturing, vol. 33, p. 101119, May 2020, doi: 10.1016/j.addma.2020.101119. 106 [158]P.-Y. Lin, F.-C. Shen, K.-T. Wu, S.-J. Hwang, and H.-H. Lee, “Process optimization for directed energy deposition of SS316L components,” Int J Adv Manuf Technol, vol. 111, no. 5, pp. 1387–1400, Nov. 2020, doi: 10.1007/s00170-020-06113-z. [159]Y. Nogdhe, A. K. Rai, and M. Manjaiah, “Numerical study of powder flow characteristics of coaxial nozzle direct energy deposition at different operating and design conditions,” Prog Addit Manuf, May 2024, doi: 10.1007/s40964-024-00652-7. [160]W. Jeong, Y.-S. Kwon, and D. Kim, “Three-dimensional printing of tungsten structures by directed energy deposition,” Materials and Manufacturing Processes, vol. 34, no. 9, pp. 986–992, Jul. 2019, doi: 10.1080/10426914.2019.1594253. [161]X. Zhang, W. Cui, and F. Liou, “Voxel-Based Geometry Reconstruction for Repairing and Remanufacturing of Metallic Components Via Additive Manufacturing,” Int. J. of Precis. Eng. and Manuf.-Green Tech., vol. 8, no. 6, pp. 1663–1686, Nov. 2021, doi: 10.1007/s40684-020-00291-7.
Thesis/Dissertation
eng
Mechanical Engineering
Vancouver : University of British Columbia Library
Attribution-NonCommercial-NoDerivatives 4.0 International
Graduate
Benoit, Michael J.
Tobber, Lisa
Performance and repair applications of 18Ni-300 maraging steel printed with micro-spot laser directed energy deposition
Text