Applied Science, Faculty of
Engineering, School of (Okanagan)
DSpace
UBCO
Wall, Andrew
Master of Applied Science - MASc
University of British Columbia
Additive manufacturing (AM) is an emerging manufacturing technique in which parts are ‘printed’ one layer at a time, which presents many unique opportunities for innovative and sustainable manufacturing, but also presents many challenges. By example, solidification cracking (SC) is a metallurgical defect that occurs during the solidification of alloys in processes such as welding, casting, and AM. SC can cause significant mechanical deterioration in parts in which it occurs, and SC remains a significant hurdle for AM processes. Research into the causes, effects, and mitigation strategies of SC in AM is ongoing, however strategies of analysing and quantifying SC is inconsistent across studies. This thesis presents a novel test for SC susceptibility in AM processes. A benchmark specimen has been produced in order to assess the viability of a self-restraint test that can promote cracking in a repeatable and quantifiable manner. The design aims to be simple, yet robust, in order to be applicable to a range of AM processes. Samples were printed by laser directed energy deposition using Inconel 625 powder for repeatability analysis, and a sample was produced using Inconel 718 powder as a comparative analysis to assess the ability to distinguish SC susceptibility between material compositions. It was found that the initial design was capable of producing cracks with repeatable results, and cracks formed in the Inconel 718 sample were less severe than the cracks produced in the Inconel 625 samples. However, the cracks were determined to be ductile fractures that initiated in the solid-state rather than SC. A redesign of the specimen is presented to address the issues facing the initial design. Changes to dimensions were made to improve printability and reduce stress concentrations in order to reduce the likelihood of solid-state cracking in favour of promoting SC initiation modes. The results provide a foundation for standardisation of SC in AM. Standardised testing is a necessary step for widespread adoption of AM, and this first step may be used for further test development as the presented design may also serve as a basis for non-SC fracture modes, such as the solid-state cracking observed in the initial samples.
ON THE DEVELOPMENT OF A NOVEL SOLIDIFICATION CRACK TEST FOR ADDITIVE MANUFACTURING by Andrew Wall A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE in THE COLLEGE OF GRADUATE STUDIES (Mechanical Engineering) THE UNIVERSITY OF BRITISH COLUMBIA (Okanagan) September 2023 © Andrew Wall, 2023ii The following individuals certify that they have read, and recommend to the College of Graduate Studies for acceptance, a thesis entitled: One the Development of a Novel Solidification Crack Test for Additive Manufacturing submitted by Andrew Wall in partial fulfillment of the requirements of the degree of Master of Applied Science. Dr. Michael Benoit, School of Engineering Supervisor Dr. Mohammad Arjmand, School of Engineering Supervisory Committee Member Dr. Lukas Bichler, School of Engineering Supervisory Committee Member Dr. Shahria Alam, School of Engineering University Examiner Dr. Zheng Liu, School of Engineering Neutral Chair iii Abstract Additive manufacturing (AM) is an emerging manufacturing technique in which parts are ‘printed’ one layer at a time, which presents many unique opportunities for innovative and sustainable manufacturing, but also presents many challenges. By example, solidification cracking (SC) is a metallurgical defect that occurs during the solidification of alloys in processes such as welding, casting, and AM. SC can cause significant mechanical deterioration in parts in which it occurs, and SC remains a significant hurdle for AM processes. Research into the causes, effects, and mitigation strategies of SC in AM is ongoing, however strategies of analysing and quantifying SC is inconsistent across studies. This thesis presents a novel test for SC susceptibility in AM processes. A benchmark specimen has been produced in order to assess the viability of a self-restraint test that can promote cracking in a repeatable and quantifiable manner. The design aims to be simple, yet robust, in order to be applicable to a range of AM processes. Samples were printed by laser directed energy deposition using Inconel 625 powder for repeatability analysis, and a sample was produced using Inconel 718 powder as a comparative analysis to assess the ability to distinguish SC susceptibility between material compositions. It was found that the initial design was capable of producing cracks with repeatable results, and cracks formed in the Inconel 718 sample were less severe than the cracks produced in the Inconel 625 samples. However, the cracks were determined to be ductile fractures that initiated in the solid-state rather than SC. A redesign of the specimen is presented to address the issues facing the initial design. Changes to dimensions were made to improve printability and reduce stress concentrations in order to reduce the likelihood of solid-state cracking in favour of promoting SC initiation modes. The results provide a foundation for standardisation of SC in AM. Standardised testing is a necessary step for widespread adoption of AM, and this first step may be used for further test development as the presented design may also serve as a basis for non-SC fracture modes, such as the solid-state cracking observed in the initial samples. iv Lay Summary This thesis attempts to address a knowledge gap in standardisation of defect formation in additive manufacturing, also known as 3D printing. In particular, a crack phenomenon called solidification cracking severely deteriorates the mechanical properties of many important alloys after 3D printing. Currently, research into the formation of solidification cracks in 3D printed parts is conducted using a wide range of methods, and this makes the comparison of results difficult. This research presents the development of a test that can be used as a foundation for standard methods to test for cracking. Ultimately, the design that was developed was able to produce cracks at pre-defined locations, which is important for standardization, however, detailed analysis revealed that the cracks were formed in the solid metal, not during solidification. Nevertheless, this design can serve as a first step towards the future development of tests and help facilitate the adoption of additive manufacturing by industry. v Preface The chapters of this thesis will be presented in a manuscript-based format and contain content from manuscripts that are published, submitted, and/or in preparation. Chapter 2 contains a literature review of existing solidification crack (SC) tests for casting and welding, as well as a review of the current state of SC research for additive manufacturing (AM) processes and a direction for future study. Chapter 4 presents the first iteration of a SC benchmark specimen for AM, including an analysis of the effectiveness of the design and the limitations and issues of this iteration. Proposed design alterations are discussed in the conclusion of this chapter based on the findings of this research. Chapter 5 presents a revised benchmark specimen design that aims to address the limitations of the previous iteration, as well as continuing to discuss feasibility of this new design. The findings are supported by microstructural characterization, numerical simulations, and feasibility assessments based on design for AM. Each chapter is a modified version of its respective manuscript, in which it has been adjusted to improve cohesion between chapters and provide a clear overarching research theme that is presented fully by this thesis. Modifications to the manuscripts attempt to reduce repeated information throughout this thesis and clearly link findings between studies. The research within this thesis was performed using funding obtained by Prof. Michael Benoit under the NSERC Discovery Grant program (RGPIN-2021-02892) grant. All research was conducted primarily by me with support from a number of individuals and groups. As this thesis contains work from manuscripts submitted for review, the research within was performed with contributions from my co-authors as well as non-author contributors. The contributions from these individuals are as follows: Prof. Michael Benoit: Supervised all work conducted as part of the completion of this project. Assisted with technical and editorial feedback for the duration of the research. Dr. Sudip Shrestha: Trained and assisted with data collection and analysis utilizing scanning electron microscopy (SEM) and associated techniques in Chapter 4. Dr. Matthew Brown: Performed X-ray diffraction (XRD) and assisted with analysis of XRD results in Chapter 4. Dr. Tony Dong: Trained and assisted with the use of Simufact Welding finite element analysis software and provided feedback on the results in Chapter 4. Ms. Emma Pugsley: Fabrication (printing) of benchmark samples and provided feedback on benchmark design presented in Chapter 4, and provided training for use of the LAWS 250 used in fabrication of the revised design in Chapter 5. As of the submission date of this thesis, the status of each manuscript is as follows: vi Chapter 2 has been published in the Journal of Materials Processing Technology under the name “A Review of Existing Solidification Crack Tests and Analysis of Their Transferability to Additive Manufacturing”. Chapter 4 is currently under review for submission to the journal Additive Manufacturing under the name “On the Development of a Novel Benchmark Design for Crack Quantification in Additive Manufacturing”. Chapter 5 is in progress, and preparation of the detailed manuscript is pending further results. Permission has been granted for the use of these manuscripts in this thesis under the condition that this thesis is not made commercially available. vii Table of Contents Abstract .................................................................................................................................................. iii Lay Summary .......................................................................................................................................... iv Preface .................................................................................................................................................... v Table of Contents .................................................................................................................................. vii List of Tables ........................................................................................................................................... x List of Figures ......................................................................................................................................... xi List of Abbreviations ............................................................................................................................ xiii Acknowledgements ............................................................................................................................... xv Dedication ............................................................................................................................................ xvii 1.0 Introduction ...................................................................................................................................... 1 1.1 Background ............................................................................................................................. 1 2.0 Literature Review .............................................................................................................................. 4 2.1 Introduction ............................................................................................................................ 4 2.1.1 Solidification cracking ..................................................................................................... 6 2.1.2 Testing for solidification cracking susceptibility ............................................................. 8 2.1.3 Special considerations for solidification cracking in casting and welding ...................... 9 2.1.4 Special considerations and challenges for solidification cracking in AM ..................... 10 2.2 Review of existing cracking tests .......................................................................................... 13 2.2.1 Solidification cracking tests in casting .......................................................................... 13 2.2.3 Solidification cracking tests in welding ......................................................................... 21 2.2.4 Existing solidification crack research in additive manufacturing ................................. 32 2.2.5 Existing additive manufacturing benchmark specimens .............................................. 36 2.3 Discussion ............................................................................................................................. 37 2.3.1 Test criteria ................................................................................................................... 37 2.3.2 Assessment of existing tests ......................................................................................... 40 viii 2.4 Conclusions and future outlook ............................................................................................ 42 3.0 Objectives........................................................................................................................................ 44 4.0 Initial Crack Benchmark Specimen .................................................................................................. 45 4.1 Introduction .......................................................................................................................... 45 4.2 Materials and methods ......................................................................................................... 45 4.2.1 Benchmark design ........................................................................................................ 45 4.2.2 Materials ....................................................................................................................... 46 4.2.3 Benchmark fabrication ................................................................................................. 48 4.2.4 Microstructural characterization and analysis ............................................................. 49 4.2.5 Simulations ................................................................................................................... 49 4.3 Results ................................................................................................................................... 50 4.3.1 Crack quantification ...................................................................................................... 50 4.3.2 Microstructure analysis ................................................................................................ 54 4.3.3 Inconel 718 benchmark analysis .................................................................................. 59 4.3.4 Solidification simulations .............................................................................................. 60 4.3.5 Thermomechanical simulations of L-DED ..................................................................... 61 4.4 Discussion ............................................................................................................................. 64 4.4.1 Effectiveness of benchmark design .............................................................................. 64 4.4.2 Crack mechanism .......................................................................................................... 66 4.5 Conclusions ........................................................................................................................... 67 5.0 Revised Cracking Benchmark Design .............................................................................................. 69 5.1 Introduction .......................................................................................................................... 69 5.2 Experimental methods ......................................................................................................... 69 5.2.1 Design improvements ................................................................................................... 69 5.2.2 Fabrication .................................................................................................................... 70 5.3 Results and discussion .......................................................................................................... 71 ix 5.4 Conclusion ....................................................................................................................... 75 6.0 Discussion ........................................................................................................................................ 76 6.1 Crack formation and mechanism .......................................................................................... 76 6.2 Quantification and standardisation ...................................................................................... 77 7.0 Conclusions and Future Research ................................................................................................... 80 7.1 Summary and conclusions .................................................................................................... 80 7.2 Limitations of the current study ........................................................................................... 81 7.3 Future research ..................................................................................................................... 82 References ............................................................................................................................................ 84 x List of Tables Table 1. Examples of AM subprocesses commonly used for metals. ..................................................... 2 Table 2. Summary of described casting tests. ....................................................................................... 20 Table 3. Summary of described welding tests ....................................................................................... 32 Table 4. Additional considerations when designing a solidification crack test for AM. ....................... 40 Table 5. Chemical composition of powders as measured by ICP-OES (wt%) ..................................... 47 Table 6. Laser power variation between layers ..................................................................................... 48 Table 7. L-DED process parameters ..................................................................................................... 49 Table 8. Crack quantification at the apexes seen in Figure 25 for IN625 replicate 1. .......................... 54 Table 9. Crack quantification at the apexes seen in Figure 25 for IN625 replicate 2. .......................... 54 Table 10. L-DED process parameters ................................................................................................... 71 xi List of Figures Figure 1. Diagram depicting typical locations for various fracture mechanisms. ................................... 5 Figure 2. (a) Hot cracking in a weld bead showing (b) the dendritic surface of a solidification crack and (c) the smooth surface of a liquation crack. ................................................................................................. 6 Figure 3. Diagram highlighting the factors and regions susceptible to solidification cracking. ............. 7 Figure 4. Spatial variation in melt pool restraint throughout printed part............................................. 11 Figure 5. Critical Diameter Method as used by Zhou et al. .................................................................. 14 Figure 6. Mould Specifications for CRC test. ....................................................................................... 16 Figure 7. Constrained Ring Casting test mould cross-section. ............................................................ 17 Figure 8. CAST test mould configuration. ............................................................................................ 18 Figure 9. Schematic of the Hot Tearing Apparatus developed by Li et al. ........................................... 20 Figure 10. Houldcroft test specimen fundamental design. .................................................................... 22 Figure 11.Double fillet weld test specimen design specifications. ....................................................... 24 Figure 12. Cast Pin Tear Test based on the levitation design. ............................................................... 26 Figure 14. Varestraint Test. ................................................................................................................... 29 Figure 15. Murex Hot Cracking Test Schematic.. ................................................................................. 30 Figure 16.The HDR Test. ...................................................................................................................... 31 Figure 17. Sample of different crack quantification methods. .............................................................. 35 Figure 19. Mechanical benchmarks as visualised by Mahesh. ............................................................. 36 Figure 20. (A) Benchmark proposed by Iuliano et al and (B) benchmark proposed by Lart. ............... 37 Figure 21. Design schematic for the benchmark specimen with a 10mm thickness. ............................ 46 Figure 22. SEM images of the (a, c) as-received powder morphology and (b,d) the particle cross-sections for (a, b) IN625 and (c, d) IN718 powders. .......................................................................................... 47 Figure 23. Scan strategy of the printed samples where n = the first layer. ........................................... 48 Figure 24. Printed samples of the benchmark specimen for both IN625 and IN718. ........................... 52 Figure 25. Optical microscopy images of cracks at the intersections of legs in both of the IN625 samples. .............................................................................................................................................................. 53 Figure 26. SEM-SE images of various locations along the crack in IN625 replicate 1, apex 2 as seen in Figure 25. .............................................................................................................................................. 55 Figure 27. SEM-SE images of various locations along the crack in IN625 replicate 1, apex 6, as seen in Figure 25. .............................................................................................................................................. 55 Figure 28. (a, d) SEM-SE images, (b, e) IPF-Z grain orientation maps, and (c, f) grain boundary maps for IN625 replicate 1 at (a-c) apex 2 and (d-f) apex 6. ............................................................................... 56 xii Figure 29. (a, c) SEM-SE images and (b, d) IPF-Z maps of the region surrounding the hairline fractures at (a, b) apex 2 and (c, d) apex 6 in IN625 replicate 1. ............................................................................. 57 Figure 30. (a) SEM-SE image and (b-e) corresponding EDS maps of a portion of the transgranular fracture for the crack observed in IN625 replicate 1 at apex 2 in Figure 25. ..................................................... 58 Figure 31. (a) SEM-SE image and (b-e) corresponding EDS maps of the tip of the hairline fracture at apex 6 in IN625 replicate 1 in Figure 25 . ..................................................................................................... 58 Figure 32. XRD analysis of IN625 replicate 1 sample. ........................................................................ 59 Figure 33. Optical microscopy of cracks located in samples produced using IN718. .......................... 60 Figure 34. Thermocalc predicted solidification behaviour of IN625 and IN718. ................................. 61 Figure 35: Simulated (a) temperature, (b) equivalent stress, and (c) plastic strain at the completion of the layer that connects each apex. Higher levels of stress and strain can be seen along the length of each apex. .............................................................................................................................................................. 62 Figure 36. Simufact Welding simulated temperature and equivalent stress vs. print time of (a) apex 2 and (b) apex 6. ............................................................................................................................................. 64 Figure 37. Design schematic of a potential redesign with a thickness of 10mm. ................................. 70 Figure 38. Initial prints of the redesigned benchmark. ......................................................................... 72 Figure 39. (a) Demonstration of slicing issues along the height of the sample due to increasing leg thicknessesalong the length of the benchmark design. (b) Although a previous layer has successful overlap between perimeter tracks, (c) the subsequent layer has insufficient overlap. ....................................... 73 Figure 40. Optical images of each apex in replicates 1 and 2 of the redesigned benchmark. ............... 74 Figure 41. High magnification optical images of cracks present in (a) replicate 1 apex 2, (b) replicate 1 apex 3, (c) replicate 2 apex 1, and (d) replicate 2 apex 3. ............................................................................. 75 xiii List of Abbreviations AM Additive manufacturing CAD Computer-aided design CDM Critical Diameter Method CPPT Cast pin tear test CRC Constrained Rod Casting CRCM Constrained Rod Casting Modified CSC Crack susceptibility coefficient CTW Controlled tensile weldability test DDC Ductility dip cracking DED Directed energy deposition EBM Electron beam melting EBSD Electron backscatter diffraction EBW Electron beam welding ECRC Enhanced constrained rod casting EDS Energy dispersive spectrometry FDM Fused deposition modelling FE Finite element FEA Finite element analysis GDP Gross Domestic Product GTA Gas tungsten arc GTAW Gas tungsten arc welding HAZ Heat affected zone HDR Hot deformation rate test HTS Hot tearing sensitivity ICP-OEF Inductively coupled plasma optical emission spectroscopy IN625 Inconel 625 IN718 Inconel 718 IPF Inverse pole figure LAGB Low angle grain boundary LBW Laser beam welding LC Liquation cracking LVDT Linear Variable Differential Transformer PBF Powder bed fusion PMZ Partially melted zone PVR Programmable deformation crack test (Programmierter-Verformungs-Risstest) RDG Rappaz-Drezet-Gremaud model SC Solidification cracking SCS Solidification cracking susceptibility SE Secondary electron SEM Scanning electron microscope SKK Suyitno-Kool-Katgerman model SLM Selective laser melting SLS Selective laser sintering xiv TB Terrabyte TIG Tungsten inert gas WAAM Wire arc additive manufacturing xv Acknowledgements I would like to begin by thanking my supervisor, Dr. Michael J. Benoit, for his immense support, both academic and personal. His kind, supportive guidance granted me the confidence to pursue things I had never considered and his assistance in navigating the many challenges of graduate study as an international student afforded me the comfort to work at the best of my ability. I have heard many stories of supervisors who follow a hands-off approach and the issues that come with that philosophy, but the constant feedback and reassurance from Dr. Benoit assured me that I was always on the right track and consequently I have developed confidence in my own skills and abilities and for that I am forever grateful. To my committee members, Dr. Lukas Bichler and Dr. Mohammed Arjmad, I am thankful for your input and professional insight into the many new fields of study I undertook during this research, as well as the freedom afforded with the shared experiences and space in our research lab. Assistance from your lab members in teaching me the use of equipment, methods of analysis, and general tips for research aided greatly in the development of this thesis. I am thankful to you and your students who have provided their input throughout my research. To my fellow lab members, with whom I spent many days of frustration and excitement, I am grateful for not only the academic input, but the friendship you provided. Taha Waqar, Chris Paul, Zahra Khodamoradi, Seyedmohammad Tabaie, Josip Vincic, Alistair Jones, and Matthew Koebel, you have all helped me keep my morale up during times of doubt, and provided input that has covered areas I would have never considered. You are all deserving of praise for your own work, and the thanks I have for your assistance towards my research cannot be understated and I will always be thankful to each and every one of you. I would like to make special mention to Taha Waqar, an international student like myself, who joined me on this journey at the same time and who I have shared each step of the way with. Overcoming the challenges facing us in a new country, in a new lab, without friends or family close by, was a journey which has cemented a unique friendship between us and his guidance as someone who has experienced graduate study before has been essential in helping me maintain optimism and focus throughout my studies. To Taha, thank you for your friendship over these past two years and I will be keeping up with your progress throughout the future of your studies. I would also like to mention Chris Paul and his friendship and support throughout my studies. Always willing to assist with my research troubles, Chris was another ear that I could voice my thoughts and feelings to, and I am grateful for his friendship over these few years. The times we spent as a group will remain fond memories regardless of where the future takes us, and I cannot express my thanks to each of you enough. The shared frustration, laughs, learning, and celebrations have made this experience an incredibly enjoyable one. xvi To my parents, Steven and Diana, your constant support and reassurance over the past two years, and the many years before then, have allowed me to pursue these dreams without fear of failure and knowing you are behind me every step of the way is impactful beyond words. I can never thank you as much as you deserve, but I hope you can understand even a fraction of the gratitude I hold for the both of you. Dad, thank you for your curiosity which I inherited, and Mum, thank you for teaching me patience and understanding. To Jennifer Sun, your presence in my life during this time has been a blessing and I will forever be thankful for all you have done for me and taught me. You witnessed the highest highs and the lowest lows, and did more than can be asked of anyone, and for that I cannot express the impact you had or the thanks I have for it. To my grandmothers, thank you for checking up on me and I’ll be home soon, to my brother Lachlan, thankyou for being a role model for me all these years, and to all my friends back home, thank you for the comfort of knowing that I’ll always have people supporting me no matter where I am. xvii Dedication Dedicated to those who supported this thesis, named or unnamed, and to all who seek only education. 1 1.0 Introduction 1.1 Background The manufacturing sector is vital to the Canadian economy, contributing 10% of the country’s GDP and providing 1.7M jobs [1]. Moreover, the manufacturing sector directly impacts several other sectors and industries including: transportation (automotive, aerospace, shipping), mining and minerals, oil and gas, textiles, etc. However, continual innovation is required in order to remain competitive in a global marketplace. One example of innovation in the manufacturing sector is the recent interest in additive manufacturing (AM) processes. AM, also referred to as 3D printing, encompasses several related technologies in which part geometries are fabricated by the selective addition of material rather than subtraction from a bulk piece of material as in conventional processes such as milling. In AM, material is ‘printed’ layer-by-layer to form a part through the fusion of the deposited material onto existing layers which reduces wasted or scrap material and allows for unique geometric possibilities, such as internal cavities or complicated lattice networks that would be otherwise challenging or outright impossible for traditional manufacturing processes. Additive manufacturing can be sub-divided into several processes or categories. In terms of AM of metals, the two main sub-categories are fusion-based or solid-state AM. All fusion-based AM processes utilize a heat source to locally melt the metallic feedstock and fuse this material with previously printed material. Fusion-based techniques can be further sub-categorized as either ‘powder bed fusion’ (PBF) processes, in which a thin layer of powder is evenly spread over a platform and a heat source is used to locally fuse powders in a given layer, and directed energy deposition (DED) processes, in which a focused heat source is directed onto the substrate or previously printed material and the feedstock material is directed into the energy source to cause it to melt and fuse with the underlying material. These processes can be further sub-categorized depending on the heat source which is used (i.e., laser beam, electron beam, electric arc) and feedstock type (i.e., powder or wire). Within the fusion-based AM processes, DED generally exhibits greater deposition rates and larger build volumes, but less geometric complexity and precision compared to PBF processes. However, DED processes can also be used to repair or refurbish existing parts [2] and can be used for enhancing parts through surface treatments (e.g., laser cladding) [3]. The focus of the research contained within this thesis is on laser-DED (L-DED), but the results are intended to be transferrable to other fusion-based AM processes. Examples of AM sub-process categories which are commonly used for metals can be in seen Table 1. 2 Table 1. Examples of AM subprocesses commonly used for metals. Process Subprocess Heat source Feedstock type Directed Energy Deposition (DED) Laser-DED (L-DED) Laser Wire or powder Electron-DED (E-DED) Electron beam Wire or powder Wire-arc Additive Manufacturing (WAAM) Plasma Arc Wire Powder Bed Fusion (PBF) Laser Beam PBF (LB PBF) Laser Powder Electron Beam Melting (EBM) Electron beam Powder Sintering (solid state) Binder Jetting Liquid bonding agents Powder Fusion Deposition Modelling (FDM) Conductive heating Combined binder/powder filament Over the past decade, AM has rapidly gained attention from both researchers and industry professionals who look to AM as a technology of great potential, especially within the context of Industry 4.0 [4], [5]. As a method of rapid prototyping, AM can accomplish goals that had previously been impossible such as the production of prototypes without tooling or machine changes and high levels of per-part customization, improving supply and manufacturing costs for low unit quantity parts needed in industries such as healthcare [6], [7], aerospace and automotive [8], and supply chain management [9], [10]. Other potential benefits of AM include: reduction of factory footprints due to consolidation of processes into a single step (AM); reduction of warehouses and inventory due to ‘on demand’ printing of parts, including the ability to easily customize designs; reduced time to market through rapid prototyping; the ability to tailor local microstructures and properties through parameter selection and/or compositionally graded materials; the ability to repair or refurbish existing parts; and the ability to perform surface treatments to enhance the performance of existing parts [11]. These benefits have led to high rates of ongoing and projected growth for AM technology and industries, with the AM market size forecast to reach $35.2 billion USD by 2027 [12]. This rapid growth has led to a great deal of interest, and the development of AM technologies and their implementation has received substantial funding from private and government sources [13], [14]. Despite this growing interest in AM, there are still large gaps of knowledge and standardised quality 3 assurance methods that hinder adoption of AM for industrial applications [15]. By example, a poor supply of highly skilled workers, lacking technical literacy within the current workforce, and high capital costs are logistical issues that has prevented industry from fully adopting AM technologies. However, there are also several technical issues that have also limited adoption, such as small selection or availability of ‘printable’ alloys, low productivity levels (i.e., slow print speeds), and poor print quality (i.e., inclusion of defects in the printed material). One defect which is particularly problematic for AM is solidification cracking (SC). As the name suggests, this defect occurs during solidification of a metal or alloy, which results in a crack in the solid material and a corresponding substantial decrease in the mechanical properties of the final part. The SC phenomenon is known to negatively affect many important engineering alloys (e.g., alloys of aluminum, magnesium, nickel, and steels) and occurs in traditional solidification-based manufacturing processes such as welding [16] and casting [17]. As fusion-based AM processes involve many individual melting and solidification events throughout printing the part geometry, SC continues to be an issue for these processes. SC is a severe and permanent defect that occurs in the fusion zone of a solidifying metal where solid and liquid phases interact at the solid-liquid boundary. As the solid phases grow dendritically and these dendrites begin to interact with one another, the flow of liquid phases in between them decreases and is unable to fill cavities that may form between the solid phases [18]. If these cavities are then subjected to sufficient stress, such as stress imposed by solidification shrinkage and thermal contraction, they can propagate in the form of a crack and cause a reduction in mechanical performance [19]. Analytical models of SC susceptibility have been developed to predict the likelihood of a crack occurring for a given set of parameters [20]–[23]. Moreover, many SC susceptibility tests have been developed for various casting and welding processes [24]–[27]. These methods serve to provide a reliable and repeatable method of SC susceptibility comparison between processes, process parameters, and materials and allow for the development and comparison of SC reduction methods. Despite this extensive background of testing for welding and casting, no such test exists for AM, making comparison of results across studies difficult, as different test procedures and crack quantification methods are hard to compare against one another to determine significant trends. Therefore, the research presented in this thesis represents the first attempt to develop a novel test or benchmarking specimen that can be used to systematically quantify SC in AM processes and, ideally, accelerate the industry adoption of AM technologies. 4 2.0 Literature Review1 2.1 Introduction Solidification cracking, also known as hot cracking, hot shortness, hot tearing, or centreline cracking, is an important phenomenon that occurs in many common engineering alloys (e.g., alloys of aluminium, magnesium, nickel, stainless steels) in solidification-based manufacturing processes such as casting and welding as cracks can severely impair the mechanical properties and, consequently, the structural integrity of the manufactured part. As fusion-based metal AM is effectively the fabrication of parts by the rapid solidification of hundreds to thousands of overlapping melt pools arranged in layers, it is not surprising that solidification cracking continues to be an issue for AM. If AM technologies are to gain widespread adoption in industrial applications, a strong understanding of these cracks is needed, including quantifiable and systematic methods to evaluate crack susceptibility. Hot cracks can be further defined by where they form. In the case of weld cracks, solidification cracks (SC) occur during solidification from the molten state, liquation cracks (D) are present in the partially melted zone (PMZ) of the filler or base metals where metal is heated into the mushy zone, and ductility dip cracks (DDC) are present in the solid metal (either solidified weld metal or base metal) where ductility drops along the solid material, such as in the heat-affected zone (HAZ) and is heavily impacted by the cooling rate [28]–[32]. Figure 1 depicts the general location of these crack initiation modes. Hot cracking and solidification cracking are often used interchangeably, but hot cracking can also include liquation cracks. The focus of this thesis will be on solidification cracks. Solidification cracking in welding and casting has been extensively studied (e.g. [16], [19], [28]–[30], [33]–[39]), however additive manufacturing has unique processing conditions and, as such, new tests will need to be developed in order to accurately evaluate solidification cracking susceptibility and bring AM to the same level of standardization as conventional manufacturing techniques that is necessary to accelerate industry adoption. 1 The contents of this chapter have been adapted with permission from Elsevier from the original publication: A. Wall, M.J. Benoit (2023), “A Review of Existing Solidification Crack Tests and Analysis of Their Transferability to Additive Manufacturing”, Journal of Materials Processing Technology (accepted). 5 Figure 1. Diagram depicting typical locations for various fracture mechanisms. 6 Figure 2. (a) Hot cracking in a weld bead showing (b) the dendritic surface of a solidification crack and (c) the smooth surface of a liquation crack. Reproduced with permission from [40] 2.1.1 Solidification cracking Solidification cracking is a metallurgical defect that can occur in the late stages of solidification. With the exception of eutectic compositions, all alloys solidify over a range of temperatures, during which it is composed of both solid grains and liquid and is referred to as the mushy zone [19], [33]. As the temperature decreases and the solid phase fraction (fs) increases, the solid grains grow and begin to interact with each other at the coherency point (T0, fs,0) and form a solid skeleton [16], [19], [33]–[35]. At this point, the flow 7 of the remaining liquid becomes interdendritic, and the permeability of the mushy zone decreases with further increases in fs, leading to pockets of isolated liquid [41], [42]. The solid grains continue to grow until the coalescence point (Tco, fs,co) is reached, at which point the grains are sufficiently connected that they form a continuous network and behave as a solid. Throughout solidification, the semi-solid material will experience tensile loading due to spatial variations in shrinkage and contraction and any mechanical restraint that is placed on the material (e.g., casting mould wall, welding fixtures). A cavity will form in the mushy zone if the flow of the remaining interdendritic liquid is insufficient to compensate for solidification shrinkage and separation of grains caused by the tensile deformation [43]. If the resultant cavity is greater than a critical size determined by fracture mechanics, it can propagate through the mushy zone as a crack [22]. It is well known that solidification cracking is related to many material and microstructure features (e.g., solidification temperature range, grain shape and size, amount and distribution of liquid, etc.) and process conditions (e.g., cooling rate, degree of restraint, etc.) [44]. Figure 3 visualises the initiation of a solidification crack. Figure 3. Diagram highlighting the factors and regions susceptible to solidification cracking. 8 Several analytical models and numerical indices have been developed to predict the susceptibility of an alloy composition or process condition to solidification cracking. Early studies, such as that by Clyne and Davies [20], defined a crack susceptibility coefficient (CSC) simply as the ratio of time available for stress relaxation to the time in the vulnerable temperature range where cracks can propagate. More physically-based models that account for a number of metallurgical and process factors have since been developed, such as the commonly used Rappaz-Drezet-Gremaud (RDG) model [21] and the Suyitno-Kool-Katgerman (SKK) model [22]. Another important aspect of these models is that they result in the determination of a critical value for a specific parameter which, if exceeded, will lead to cavity formation and subsequent cracking. For example, the RDG model is used to determine the maximum deformation rate that can be sustained in the mushy zone before a cavity occurs at the root of the dendrites, whereas the SKK model compares the mushy zone stress and the size of shrinkage cavities to determine if a crack will occur using the fracture mechanics. However, these models require a large number of material property and process condition inputs, which may not be well known, particularly for the rapid cooling rates experienced in AM. Recently, numerical indices have been developed to assess the solidification crack susceptibility of an alloy using only the T vs. fs curve of the material as an input, such as the indices developed by Kou [43], [45]–[47] and Easton et al. [23]. While these indices have gained popularity in recent years due to their relative simplicity, they only provide a measure of the relative crack susceptibility between alloys of similar composition and do not prescribe a critical value for crack formation. Moreover, the indices require the selection of the coherency and coalescence points, which are often not precisely known but can significantly influence the predicted crack susceptibility of an alloy [48]. 2.1.2 Testing for solidification cracking susceptibility Due to the potential for unexpected structural failure of a component as a result of impaired mechanical properties caused by cracking, it is necessary to be able to evaluate the susceptibility of an alloy composition and process conditions to solidification cracking reliably and quantifiably. As such, over 140 solidification cracking susceptibility tests have been developed for casting and welding [16]. While hot cracking has been studied for decades, e.g. [30], [36]–[39], no existing criteria can definitively state if a crack will occur during solidification, though they can provide very strong indication of crack susceptibility, albeit with the requirement of extensive material and process knowledge [34]. The RDG and SKK models remain valuable methods of crack susceptibility analysis, as if the required input data is available, they can provide a high level of accuracy of whether or not a crack will occur. However, they do not easily predict the severity of a crack or where a crack will occur for a given geometry, unless coupled with finite element analysis (FEA). These issues must be dealt with through practical tests which can not only show the severity of the crack, but can be performed without much background knowledge on the material properties. Due to practical 9 tests not requiring extensive and accurate background data, they are more robust and less susceptible to errors in assumptions made during the experiment. This is partly why so many tests have been developed, as each one attempts to evaluate a specific criterion that is often unique to the manufacturing process under consideration [16], [37], [49]–[53]. Solidification cracking tests are typically divided into two main categories based on the applied load [53]: self-restraint and externally loaded. Self-restraint tests are based on the aforementioned principles of thermal contraction and solidification shrinkage in which the alloy will develop internal stresses that can cause cracking without an external load. Conversely, externally loaded tests rely on repeatable and measurable external loads, typically applied by specially designed machines. The measurable loads coupled with crack measurements provide quantitative data that can be used to assess crack susceptibility, as opposed to the more qualitative data produced by self-restraint tests that simply test for the existence or non-existence of a crack. 2.1.3 Special considerations for solidification cracking in casting and welding 2.1.3.1 Solidification cracking in casting Casting typically involves a large amount of liquid metal (compared to welding) solidifying in a mould. The thermal gradient varies throughout the liquid as it cools, and this spatial variation is affected by the metal pouring temperature and the temperature and physical properties of the mould [54]. The varied thermal gradient can result in uneven cooling conditions and, thus, directional solidification, resulting in a spatial variation in the solid fraction throughout the casting, causing the aforementioned variations in thermal contraction and solidification shrinkage that leads to the development of thermally induced stresses on the mushy zone. The magnitude of solidification cracks in a casted specimen can range from those visible only at a microstructural scale, to those that completely bisect a test specimen [51], [52]. 2.1.3.2 Solidification cracking in welding Solidification cracks in welds, like casting, can have a large variance in the magnitude of the cracks, ranging from small cracks detectable only with microstructure analysis, to large cracks that propagate down the entire length of a weld [16], [55]–[59]. Solidification cracking in welding differs from casting as the liquid weld pool is much smaller than a typical casting which usually results in faster cooling rates, and the solidifying weld pool is surrounded by the solid base metals being joined, which can result in epitaxial growth, whereas a bond between the liquid metal and the solid mould is unwanted in casting. Welding also involves a moving localised heat source which can lead to the development of residual stresses throughout the weldment that apply a load to the solidifying metal. The welding parameters affect the solidification conditions and corresponding microstructure evolution within the weld pool. For example, a faster weld speed will result in a teardrop shaped weld pool, while a slower speed results in an elliptical shape [19]. The shape of the weld pool affects the direction of columnar dendritic growth, as grains tend to grow 10 perpendicular to the melt pool boundary along their “easy growth” direction (<100> for cubic materials) [19], [60]. Grains with their easy-growth direction oriented perpendicular to the weld pool boundary will grow more easily and crowd out less-favourable grains (i.e., competitive growth), resulting in a dominant orientation of grains perpendicular to the weld boundary [19]. Therefore, the weld speed and weld pool shape affect the grain structure and in turn can affect solidification cracking in the weld. Similarly, the process conditions can affect both the thermal gradient (G) and the growth rate of the solid/liquid interface (R) and, consequently, the degree of constitutional supercooling, grain size and morphology, and crack susceptibility. However, this is common throughout all solidification processes and is not limited to welding. Solidification cracks in welding are not usually caused solely by the internal stresses induced by cooling; external loads that may be present can induce cracking as well, e.g., from fixtures or clamps on the base plate or from the mould in the case of casting processes, and this is the basis of the externally-loaded tests that will be discussed in this review. 2.1.4 Special considerations and challenges for solidification cracking in AM Solidification cracking in AM follows the same basic phenomenology of solidification cracking in casting and welding, though with several unique complications. Microstructures of materials fabricated by AM processes are often characterised by coarse columnar grains that grow epitaxially along the build direction, and often growing through multiple layers due to re-melting of the prior layer during printing [61], [62]. The sample orientation and scan pattern during printing also affects the solidification direction and, consequently, the microstructure, which means that parts that have the same geometry can have different microstructures [41], [63]. The small melt pools common throughout AM processes (e.g., <1mm for laser powder bed fusion (L-PBF) to a few mm for wire–arc AM (WAAM)) lead to high cooling rates and rapid solidification, causing unique issues such as localised elastic strains [63] and the unreliability of classic Scheil equations to determine susceptibility to solidification cracking within AM [64]. Some additional complexities for solidification cracking in AM come from the issues of complex thermal cycling and resultant loading conditions. AM imparts a unique thermal history to the printed material due to layers being heated and cooled as subsequent layers are deposited on top. For a given point in the initial layer of a part, it first solidifies and cools as the rest of the layer is printed, but as the next layer is printed on top of it, a portion of the initial layer is re-melted and solidified. During printing of subsequent layers, the material in the previous layer is reheated and cooled, with the peak reheat temperature decreasing with increasing number of layers. In addition to repeated heating and cooling with the printing of new layers, individual melt tracks can be re-melted or re-heated during printing of additional tracks in the same layer. Tracks overlap one another within a given layer as defined by the hatch spacing and, as such, the heat can be conducted from the liquid melt pool by the adjacent, previously printed track. In a similar fashion, the 11 load conditions experienced by the melt pool can vary spatially within the part and within the same layer. A schematic representation of how the degree of restraint and heat flow out of the melt pool can vary throughout a build is illustrated in Figure 4. While not shown conclusively, it can also be expected that the nature of the restraint would differ between AM processes. For example, the melt pool in L-BPF is surrounded by metal powder in the same layer, whereas other processes such as laser directed energy deposition (L-DED) or WAAM would only be restrained by the already solidified metal within its own layer and previous layers. Figure 4. Spatial variation in melt pool restraint throughout printed part. Generally speaking, solidification cracking in AM has not been studied as extensively as in welding and casting, but that can largely be attributed to the relative infancy of AM. Within AM, solidification cracking can depend on the scan speed during printing [65], which is similar to the relationship between solidification cracking and weld speed for welding processes. Cracking typically occurs parallel to the build path [66], consistent with the growth direction of the dendrites and the direction of the thermal gradient, and which is common among both welding and casting processes. AM cooling differs from casting as the molten metal in AM is in constant contact with already solidified material which provides a conductor of heat, shaping the thermal gradient according to the scan direction and build pattern, as opposed to the cooling of a cast 12 metal which cools depending on the mould dimensions [67]. The repeated thermal cycling between layers and the overlap of tracks within the same layer affects the microstructure of the material making predictions of the as-printed microstructure difficult [68], [69]. The geometry of the specimen can heavily impact the thermal history too, as the 3-dimensional structure can direct heat-flow differently depending on the surrounding material, as opposed to welding processes in which the heat source remains in a 2D space and won’t generate heat above or directly beside a previously solidified track. With the unique conditions of AM processes that could influence solidification cracking comes a need for unique tests to analyse the susceptibility of different materials and processes to solidification cracking. Existing benchmark specimens exist for AM, and will be reviewed briefly later in this article, but these tests typically focus on the geometrical constraints and mechanical properties of 3D printing [70]–[72], with no analysis of the solidification cracking susceptibility. Solidification cracking in AM has been studied before [61], [73]–[76], but the collective research suffers from a lack of standardisation and quantitative data. The existing research on solidification cracking in AM generally has somewhat consistent findings (e.g., reduction of cracking with reduced laser scan speed in L-PBF, and a reduction or elimination of cracking when inoculant particles are used to create fine, equiaxed grains [48], [77]). However, the lack of standardisation and quantification of cracks reduces the ability to compare results confidently between studies, making it difficult to determine the best crack mitigation or elimination methods. Cracks have been quantified based on their individual lengths and widths [66], as a ratio of defect surface area to total surface area [48], [77], or even using simulations and theoretical cracking criterion [78]–[80]. However, in order to compare results from a given experiment to existing literature, a robust, standardised, and quantifiable measure of cracking must be developed. This is the key reason as to why a standardised benchmark specimen is required within the field, much like existing standards within welding and casting. If these test procedures could obtain some level of standardisation, then the AM industry would benefit from the ability to rank materials, processes, or machines in terms of solidification crack susceptibility, allowing informed decision-making for AM production choices. Despite the significant degradation solidification cracking can have on material properties, few attempts have been made to systematically and quantitatively study solidification cracking in AM. The objective of this chapter is to set a direction for the development of a standard test specimen that can be used to systematically and quantifiably study solidification cracking in AM. Several existing tests used in casting and welding designed to evaluate the susceptibility of a material or process parameters to solidification cracking will be reviewed. Current trends in AM benchmark specimen design will also be discussed and a list of criteria for developing an AM cracking benchmark will be outlined. Lastly, the casting and welding tests reviewed will be evaluated on their applicability to AM based on the proposed criteria. 13 2.2 Review of existing cracking tests 2.2.1 Solidification cracking tests in casting The nature of casting processes presents some challenges when developing hot cracking tests, as it is not feasible to directly apply an external load to a molten metal. Consequently, loads are usually applied in a different way for casting tests. Tests such as the Ring Crack Test [81]–[84], the Cast Pin Tear Test (CPTT) [85]–[90], and the Constrained Rod Casting (CRC) test [27], [51], [52], [54] all use the force applied by the mould as the molten alloys solidifies and contracts inside of it to evaluate cracking. As the load applied through the mould geometry cannot be directly controlled, the resultant data is qualitative. However, approaches have been taken to obtain supplementary, quantitative data for these tests. By example, Dubey [51] modified the standard CRC test to include a measurement rod to allow for thermal contraction force measurements alongside the qualitative crack measurements. It should also be noted that the mould not only affects crack susceptibility of an alloy by constraining solidification shrinkage, but also through the development of thermal gradients in the molten metal [21]. 2.2.1.1 Critical Diameter Method (CDM) The key principle of the CDM is to cast a set of equal length rods that vary in diameter as seen in Figure 5. Rods with smaller diameters are more likely to crack, as they have higher cooling rates compared to larger diameter rods and, consequently, a faster deformation (shrinkage) rate, as well as having a higher stress resulting from a smaller cross-sectional area [91]. From the test results, the crack susceptibility can be determined by identifying the largest (critical) diameter specimen that experiences solidification cracking. Smaller increments of diameter change can increase the sensitivity of the test results, but repetition of the test to find an exact diameter of cracking can be costly. Therefore, the results instead can be used qualitatively to compare the critical diameter between different alloys, pouring temperatures, mould temperatures, or other casting parameters, in order to rank them against one another in terms of solidification crack susceptibility. This test falls under the sub-category of tests known as ‘Backbone Tests’, which includes other tests such as the Constrained Rod Casting test, and has been used for materials such as 7075 [91], Mg-6Al-0.5Mn [92] and various Al-Cu alloys [93]. 14 Figure 5. Critical Diameter Method as used by Zhou et al. Image reproduced with permission from [91] 2.2.1.2 Constrained Rod Casting (CRC) The CRC test follows a similar idea to that of the CDM, except that the cast rods in the CRC are of the same diameter but different lengths, as shown schematically in Figure 6 [27]. An important aspect of the CRC test results is how results are quantified. Cao and Kou assessed crack severity at three points along the length of the rods and gave a numerical rating based on the length of the rod, the width of the crack, and the location of the crack along the rod [27]. Kamga et al. [94] developed a new Hot Tearing Sensitivity (HTS) index in order to quantitatively assess the crack severity based on the size of the crack and cast rod the crack was observed in. A numerical value of cracking is given to each crack on a scale from 0 (no cracking) to 4 (complete fracture). The HTS calculation then assesses a given alloy based on the level of cracking found within each rod length, and normalises it so that if, for a given alloy, all rods are completely fractured after casting, HTS = 1, and if no hot tears are formed in any of the bars, HTS = 0. The calculation for HTS is written as: 𝐻𝑇𝑆 = ∑ 𝑐𝑖/𝐿𝑖𝐷𝑖=𝐴∑ 4/𝐿𝑖𝐷𝑖=𝐴 eq. (1) in which Ci is the numerical value (0-4) assigned to crack severity in rod i = A, B, C, or D, and Li is the length of the rod. The normalisation of results in the HTS calculation serves to create a consistent data set between experiments. There have been multiple modifications to the traditional CRC test geometry, such as the Constrained Rod Casting Modified Horizontal test (CRCM-Horizontal) designed by Malau et al. [52], as well as the Enhanced Constrained Rod Casting (ECRC) method developed by Dubey [51]. As the name implies, casting in the CRCM-Horizontal test is performed horizontally rather than vertically, and the mould 15 is changed from the ‘harp-shaped’ design used in the traditional CRC (Figure 6) to a ‘star-shaped’ mould. This change on the mould design allows the molten metal to flow into each rod at the same time, and results in a more even heat transfer between rods; in the traditional CRC, the top and bottom rods will only absorb heat from rods on one side, whereas the middle rods will be affected by heat transfer from the rods on both sides. The ECRC test addresses the issue of molten metal flow in a different way by introducing a tapered sprue to fill the traditional harp-shape mould. This adjustment aids in filling each rod cavity at the same time, allowing for near simultaneous solidification of each rod, similar to the CRCM-Horizontal design. The ECRC design also includes a thermocouple and load cell in order to measure contraction forces during solidification. Such load cells have also been used by D’elia et al. [54] to determine the temperature at which solidification cracking occurred. The CRC test and its variations have been used for binary Mg-Al alloys [27], Al-Zn-Si and Al-Cu-Mg alloys [52], A206 and A308 alloys [51], and a number of other Al or Mg alloys [95]–[97]. The CRC and its variations exhibit the importance of understanding thermal gradients within the test geometry in order to create consistent results. This is worth considering for an AM crack specimen since the thermal gradient can change as the part is printed and thermal accumulation occurs. Another important aspect of the test that increases the consistency of results is the pre-defined points of analysis and clear definitions of crack severity. By designing a test with defined areas to assess as well as an understanding of the level of cracking, standardisation can be improved within AM hot cracking research which is lacking in the current literature. 16 Figure 6. Mould Specifications for CRC test. Reproduced with permission from [27] 2.2.1.3 Constrained Ring Casting Constrained Ring Casting, also referred to as the Ring Mould test, is among the most common casting solidification cracking tests, due in part to its simplicity and ease of implementation [52]. The mould consists of an annular cavity in between an outer ring and an inner core, as shown schematically in Figure 7. As the metal cools and solidifies it shrinks around the inner core resulting in the development of tensile stress throughout the solidifying metal and produces cracks parallel to the solidification growth (i.e., self-restraint test). The lengths of these cracks are measured and are used as the principal result. Similar to other casting crack tests, the mould can be preheated to reduce expansion of the mould and the risk of over-stressing the specimen. Cooling of the inner core can also be used to increase the cooling rate of the metal by, for example, circulating cooling water through the inner core, as used by Drezet [83]. The test has been used on Al-Zn-Mg alloys [81], Al-Cu alloys [98], and even non-metal materials such as concrete [99], showing that the simplicity of its design allows it to be widely adopted across different applications. 17 Figure 7. Constrained Ring Casting test mould cross-section. Reproduced with permission from [33] 2.2.1.4 CAST Hot Crack Apparatus A unique casting solidification crack test was developed by Instone, StJohn, and Grandfield [100] that allows for three loading conditions during solidification: free contraction, constrained contraction, and applied extension. Both free and constrained contraction are self-restraint methods whereas the applied extension is an external loading method. Two cast bars are assessed in the CAST test as shown in Figure 8; one bar is used to assess the load development during solidification (tensile bar) and the other bar is used to record temperatures. The bars are cast at the same time from a central sprue but are on opposite sides of the apparatus so that the thermal conditions are not affected by tensile loading equipment. The tensile bar is of particular interest due to the ability to test multiple load conditions with the same apparatus, and these conditions can record data on contractive loads and mushy zone strength that can be correlated with the temperature and solid fraction data obtained from the other bar. As the bar solidifies, it does so around a moveable end plate. If this end plate is not restrained, the casting is free to solidify without any hindrance from the apparatus and this is referred to as free contraction. If this moveable plate is instead restrained, then it will apply a reactive force against the contraction of the solidifying bar, and this is referred to as constrained contraction. Constrained 18 contraction is still a self-restraint process, as all induced force is caused by the result of solidification shrinkage and contraction. The last method of testing is applied extension in which the end plate moves in the opposite direction of solidification by way of a tensile loading machine. As with externally loaded tests found in welding, this tensile loading is pre-defined and, as such, numerical data can be obtained and assessed. A load cell and LVDT are present to measure the load development and contraction during solidification for all 3 loading methods. This apparatus was used on AA7075 and AA194 to determine strength development in the mushy zone, though the authors note its potential for use with foundry alloys and magnesium alloys. The study assessed cracking qualitatively using optical microscopy to validate results, however quantitative measurements could be taken using methods such as assessing total crack length or crack surface area. Figure 8. CAST test mould configuration. Image reproduced with permission from [100] 2.2.1.5 Warrington and McCartney Test The solidification cracking test developed by Warrington and McCartney [101] is another casting test which utilises thermocouples to measure cooling rates and local solidification times. This test differs from the previously discussed tests in that it utilises a crucible with molten metal and a copper chill, as opposed to a mould configuration. In this test, the molten metal is held in the crucible at a desired temperature and then the conical, water-cooled copper chill is plunged into the liquid to a given depth. Solidification of the molten 19 metal occurs in a direction approximately perpendicular to the chill surface and the resulting solid is a thin-walled conical section. A vertical stripe of colloidal graphite is painted on the surface of the cone prior to insertion, and it is at the location of this stripe that a solidification crack would form due to an induced hotspot caused by the stripe influencing cooling behaviour. The crack length at this location is used as the comparative output of this test. This test was developed for use with Al alloys and has been used for such [102], though the concept could be adapted for use with other materials. One of the advantages of this test is in its ability to direct the growth of grains and test crack susceptibility perpendicular to this growth direction. By using the geometry of the design to direct the solidification direction, the susceptibility to cracking based on a desired orientation of the grains can be obtained. The ability to control the direction of grain growth could be relevant for an AM crack specimen design since, as previously noted, AM microstructures are often characterised by coarse columnar grains and epitaxial growth. 2.2.1.6 Hot Tear Rating Apparatus While evaluating the effect of alloy composition of 7xxx Al alloys, Li et al. [103] developed a new test apparatus with a load cell and thermocouple to determine the load for a given solid fraction as shown schematically in Figure 9. The apparatus is based on a similar design developed by Eskin et al. [104] and functions in a similar way to the CAST test apparatus in that the cast metal solidifies around a screw fixed to a moveable wall which is attached to a load cell that can measure load development throughout solidification. This apparatus allows for the collection of load data and thermal measurement which can be compared to the crack width of the tested specimen. The results of the test elucidate the relationships between alloy composition, the maximum solidification load, and crack length through the load cell measurements. An important distinction between this apparatus and the CAST apparatus is that the CAST apparatus involves the solidification of 2 separate cast bars, one for load measurements and one for thermal measurements, whereas the test developed by Li et al. requires a singular cast bar. Although Li et al. used this test to study 7XXX series alloys, the test used by Eskin et al. studied 6XXX series alloys and the straightforward apparatus could be easily adopted for other materials. 20 Figure 9. Schematic of the Hot Tearing Apparatus developed by Li et al. Adapted from [103] with permission. 2.2.1.7 Casting test summary The focus on self-restraint tests in casting provides important insights into design concepts for potential AM tests, as any AM test will also likely rely on stresses induced by self-restraint during solidification. Certain methods of quantifying the qualitative results of self-restraint tests through the use of instruments such as thermocouples and load cells might be difficult to introduce into the build for some AM processes. However, other methods of quantifying the data can be used, such as the assessment of crack length and severity at predefined inspection points, as demonstrated in the CRC test. A summary of the casting hot tearing tests discussed in the section is provided in Table 2. Table 2. Summary of described casting tests. Test Applied Force Loading Loading direction relative to solidification Test Output Critical Diameter Method Self-restraint - - Minimum rod diameter for crack CRC Self-restraint - - HTS index Constrained Ring Self-restraint - - Total crack length CAST Hot Crack Apparatus Multiple Tensile Transverse Load vs fraction solid Warrington and McCartney Self-restraint - - Crack length Y. Li et al. Self-restraint - - Crack width, load, and solid fraction 21 2.2.3 Solidification cracking tests in welding Welding solidification crack tests benefit from the fact that welds are made between solid materials to which an external load can be applied with repeatable and controlled conditions. Many welding tests take advantage of this fact and test a wide variety of load conditions. Tests such as the Varestraint and Transvarestraint Tests [16], [30], [36], [50], [53], [81], [105], the Sigmajig test [16], [49], the Controlled Tensile Weldability test (CTW) [16], [106]–[108], and others are examples of externally loaded tests for cracking susceptibility. The different weld and load conditions that have been studied have resulted in a large variety of design approaches and each test possesses a unique set of benefits and limitations. Consequently, there are varying levels of adoption for each test with some tests such as the Varestraint test and its variants becoming ISO standardised methods [105], [109]. While having a controlled external load allows for quantifiable force data to be known, not every weld test is externally loaded. The Houldcroft test is a self-restraint test that is widely used [16], [105], [110]. The variety of solidification cracking tests used in welding indicates the vast number of criteria that are being assessed, which will be discussed in this section. The tests reviewed are a subset of the total number of tests developed, but they represent the most common tests, tests with a unique approach, or tests with a specific relevance to AM processes. 2.2.3.1 Self-restraint welding tests Self-restraint tests largely benefit from their simplicity compared to externally loaded tests, as they generally consist of a simple test specimen that is to be welded without an external testing jig. This usually results in an inexpensive setup and simple data, which leads to self-restraint tests being good tools for qualitatively ranking materials or welding processes. For example, crack sensitivity between alloy composition or weld process parameters can be compared through visual analysis and crack length measurements to rank them without knowing values such as the stress or strain acting on the mushy zone which induced the cracking. 2.2.3.1.1 The Houldcroft (Fishbone) Test One of the most well-known self-restraint solidification cracking tests is the Houldcroft test, which is often referred to as the fishbone test due to the specimen geometry as shown schematically in Figure 10 [25]. The test consists of a single metal plate with slits cut into both sides of the specimen, with the depth of the slit increasing with distance along the weld length. The weld direction conventionally goes from the side with the shortest slit length (i.e., highest restraint on the weld) towards the end with the longest slit length (i.e., lowest restraint on the weld), although the reverse has also been used [110]. The Houldcroft test is used to analyse the susceptibility of the weld material to crack propagation, as opposed to crack initiation. A crack 22 is initiated at the start of the weld, and crack susceptibility is determined by comparison of the final crack length after welding to the total specimen length. As the weld progresses, the restraint on the weld imposed by the slotted baseplate gradually decreases and, therefore, longer crack lengths indicate a greater crack susceptibility. The Houldcroft test does not predict whether cracking will occur in a specimen, but rather how large a crack could become if one was to occur. The Houldcroft cracking test has been used on a wide variety of materials to good results, such as mild steels [110], AA5052 aluminium [111], other Al alloys [112], as well as experimental materials such as TiNb alloys [113]. The Houldcroft test uses the geometry of the specimen to impose varying levels of stress along the weld, which could be a very useful concept for the development of an additively manufactured cracking test, as the geometry of the design will be the easiest and most flexible variable to adjust in test or specimen design. Figure 10. Houldcroft test specimen fundamental design based on [25]. 2.2.3.1.2 The Fan-Shaped Cracking Test Solidification cracks are not observed in the weld bead when crack tests that are typically used for arc welding are used for electron beam welding (EBW), due to the higher energy density and weld speed and correspondingly narrower heat affected zone (HAZ) associated with EBW. Matsuda and Nakata [37] developed an alternative Houldcroft test geometry specifically for EBW in which the specimen has a shape like an unfolded fan. The shape and dimensions of the specimen for this test are recommended by the authors for both aluminium alloys and stainless steels for EBW. 23 The output of the test is similar to the Houldcroft test in that it compares the length of a single centreline crack to the length of the specimen, with crack susceptibility being given as a crack percentage. The crack propagates due to a deflection force caused by thermal expansion and moves from a point of lower restraint to higher restraint, as the increasing width provides an increase in restraint force which counteracts the deflection force, and thus the crack will cease propagation at a critical width where the restraint force is high enough to prevent further propagation. The fan-shaped test emphasizes the importance of the heat source during welding, as the heat source can affect crack susceptibility through total heat input and travel speed, and could be an important consideration in the design of an AM crack specimen. 2.2.3.1.3 The Double Fillet Weld Test The double fillet weld test is another self-restraint test with a fairly basic geometry. The test consists of a horizontal base plate with a perpendicular plate to be welded (i.e., tee joint). As the name suggests, the test consists of two fillet welds, one on either side of the tee joint, though only one is used for crack susceptibility testing. Three tee joint dimensions and configurations are standardised in ISO 17641 for the double fillet weld test as shown in shape 1 shape 2 24 Figure 11 [109], with each one designed to impose a different degree of restraint on the weld. Shape 1 is the standard test piece, whereas shapes 2 and 3 impose higher levels of restraint, and can be used interchangeably based on access to materials. The first weld must be produced crack free through standardised welding procedures and serves to be a source of suitable restraint for the specimen, whereas the second weld is produced with 20% less volume and is used for crack susceptibility analysis [16]. The test simply evaluates for the presence of cracks (i.e., crack or no crack) and not severity. The measuring points shown in shape 3 shape 1 25 Figure 11 are the locations at which weld thickness is measured for volume control. This test is often used for testing dissimilar filler metal-base plate combinations, such as 316LN austenitic stainless steel and Inconel 800 [114], [115]. The concept of joining two separate sections of a specimen largely inhibits the transferability of this test to AM processes, as AM is not used for this purpose. However, the design of the geometry is simple and weld 1 serves as a controlled, crack free specimen with weld 2 being restrained by weld 1 for crack analysis. This is a concept that could be used in an AM specimen in order to provide some level of control into a self-restraint test that might otherwise suffer from a lack of crack stability which can inhibit quantitative analysis. shape 2 shape 3 26 Figure 11.Double fillet weld test specimen design specifications reproduced with permission from [109] 2.2.3.1.4 The Cast Pin Tear Test (CPTT) The cast pin tear test was developed in the 1960s by Hull [26] to test the susceptibility of different alloys to solidification cracking. This test is interesting in that while it is often cited in welding literature, it does shape 1 shape 2 shape 3 27 not actually involve welding in the test. Instead, the test involves casting a small volume of a metal into a copper mould (Figure 12) which affects the restraint applied on the solidifying metal by way of the mould geometry. The solidifying metal is cast into a pin shape, which is then restrained on both ends by the mould as it cools, applying stress based on the geometry and thermomechanical properties of the mould. The solidification shrinkage in the liquid metal and the thermal expansion in the copper mould due to heat extraction from the liquid results in an applied force on the two heads of the pin, promoting crack initiation in the specimen. Much like the CRC Test, the CPTT is performed with a series of pins of increasing length, with cracking increasing with increasing pin length. The pins are assessed using optical imaging to quantify crack severity based on the percentage of circumferential cracking, which is then plotted alongside the mould used, as the geometry of the mould impacts crack initiation. By comparing the crack severity between pins of different materials produced using the same mould, their crack susceptibility can be ranked. A detailed overview of the test was given by Alexandrov and Lippold [85]. While no welding process is used in CPPT, its use for crack susceptibility in welding should not be understated as there have been multiple variations developed [116], [117]. The more recent versions of this test standardise the mould geometry and remove the need for arbitrary mould numbering and instead use pin length to plot crack susceptibility. The original test developed by Hull utilised levitation heating and was focussed on stainless steels, however newer generations of the test replace this method with gas tungsten arc (GTA) melting and has been used for materials such as Ni-based filler metals [85]–[88]. The CPPT demonstrates the mechanisms that could be used to induce tensile loading on the mushy-zone in a self-restraint test, such as the use of a mould (or build plate) material with high thermal conductivity or the physical design of the mould (or build plate) [85]. 28 Figure 12. Cast Pin Tear Test based on the levitation design. Adapted with permission from [118] 2.2.3.2 Externally loaded welding tests Externally loaded tests are typically more complex compared to self-restraint tests, as they inherently involve mechanisms to provide a controllable load, but the data obtained is often more quantitative. These tests often have standardised experimental setups which aid in the consistency and reliability of data. 2.2.3.2.1 Sigmajig Test The Sigmajig test was developed by Goodwin in 1978 and is an externally loaded test that pre-applies a transverse stress to a specimen prior to welding [49]. The test specimens used during development of the test were 50 × 50mm2 square sheets of austenitic stainless steel, though the test has been used for other materials such as 5000 and 6000 series aluminium alloys [119]. The application of stress is followed by autogenous gas tungsten arc welding (GTAW) along the length of the specimen. Several repetitions of the test are performed for the same GTAW parameters and alloy composition with increasing levels of stress until cracks eventually occur. This test suffers from the requirement of multiple trials to find the critical 29 pre-stress to cause cracking, which may then need re-testing to confirm results. However, the test does provide a precise critical stress value and starting stress, as well as real-time measurement of loads (stress drops throughout the test based on the amount of yielding experienced by the specimen) allowing for quantifiable results such as applied stress vs. percent crack length or critical stress required to induce cracking. This data can be analysed to determine crack susceptibility between materials and rank them accordingly. 2.2.3.2.2 Programmable Deformation Crack Test (PVR) The PVR (Programmierter-Verformungs-Risstest) test is different from the Sigmajig test in a number of areas, with one of the most important being that samples in the PVR test undergo a linearly increasing tensile stress during welding, as opposed to a fixed stress applied before welding [120]. Moreover, the load is applied parallel to the weld direction, rather than transverse to the weld direction. Consequently, the cracks that occur in PVR tested specimens typically occur transverse to the weld direction [16], as opposed to the centreline cracks seen in many other tests such as the Sigmajig, Houldcroft, Double Fillet Weld, and Varestraint tests. The outcome of the test is the critical deformation rate at which a crack occurs that is then compared between samples. This test has been used for stainless steels as investigated by Herold et al. [120] and Inconel alloys as investigated by Slyvinsky et al. [121]. The PVR test highlights the importance of considering the direction of crack formation and growth when designing a crack test. The internal loads developed during solidification in AM will vary based on the geometry of, and location in, the printed part, and the orientation of the print will have an effect on this. The variation in load based on the location in the printed part is due to the multiple tracks in a layer, where the initial track will be constrained by the layer below it and the material deposited behind it, whereas subsequent tracks will also be constrained on the side by the tracks deposited before them, as shown in Figure 4. 2.2.3.2.3 Controlled Tensile Weldability Test (CTW) The CTW test, developed by Kromm and Kannengiesser [122], is another test that applies a tensile load during welding. However, unlike the PVR test, the weld direction can be varied relative to the load direction, as applied by Gollnow and Kannengiesser [107]. The strength of the CTW test is that it can be used for a variety of sample dimensions, materials such as Al alloys [123] and steels [106], and weld processes such as GTAW [107] and laser beam welding (LBW) [106]. A defined loading rate is applied to the specimen before or during welding through a clamping device that is actuated by servomotors, and this precise load application can be used to determine critical tensile loading rate at which cracks occur for a given specimen and weld conditions. However, due to this flexibility, it is important to carefully interpret the results, as if 30 the load, loading rate, or dimensions of a given sample vary between tests, the final results may differ and an understanding of the variables at play is important to come to an accurate conclusion. Similar to the PVR, the CTW test also highlights the importance of loading direction and crack growth direction, but unlike the PVR test, different sample sizes and weld directions are able to be employed. This additional flexibility is an important factor to consider in terms of adaptability to AM, as designing a test that is suitable for different processes in terms of its size and printability will improve its applicability among AM processes. 2.2.3.2.4 Varestraint and Transvarestraint Tests The Varestraint and Transvarestraint tests are among the most commonly used crack susceptibility tests used for welding. These tests employ a bending load during welding; the classic Varestraint test applies a bending load longitudinally along the weld direction (Figure 13), whereas the Transvarestraint test bends the specimen transversely. The tensile strain imparted during the test can be altered by adjustment of the specimen thickness or the radius of the die which the specimen is bent around, as defined by Equation 2. 𝜀𝑡𝑜𝑡𝑎𝑙 ≈ 100 𝑥 ℎ2 𝑥 𝑅𝑚 eq. (2) in which: Ɛtotal is the tensile strain on the specimen surface, h is the specimen thickness, and Rm is bending die radius. This flexibility to adjust the severity of the applied load through relatively simple means aids in the adaptability of the tests and, thus, their wide application. Both of these tests allow for quantitative comparison of results between experiments and have been standardised in ISO TR 17641:3 [109]. The analysis of the results is standardised in which an alloy’s resistance to solidification cracking is determined by quantitative analysis of the total length of the centreline crack in the Transvarestraint test and by total crack length in the Varestraint test. The Varestraint test has been applied to a wide variety of materials, ranging from steels [59], [113], [124], aluminium alloys [105], [125], [126] and nickel alloys [30], [127], including additively manufactured Inconel [128]. The Varestraint and Transvarestraint tests offer valuable insights for the design of hot cracking tests. Adapting sensitivity to applied strain through dimensional adjustments is an important consideration, as AM tests will largely rely on their dimensional constraints to impose strain and stress. When data must come from qualitative sources, being able to standardise this qualitative data will prove valuable in the consistency of results between experiments. A qualitative assessment can be used to create a quantitative analysis which helps lead towards standardisation. 31 Figure 13. Varestraint Test. Image reproduced with permission from [16]. 2.2.3.2.5 Murex Hot Cracking Test The Murex hot cracking test is similar to the Transvarestraint test, as it tests the hot cracking susceptibility by inducing a bending load during welding. However, the Murex test uses two separate specimens welded together, with one specimen being held stationary while the other specimen rotates around the longitudinal (welding) direction as shown in Figure 14. The Murex test is most often used on steels [129], [130] and though the susceptibility of steels to hot cracking can be determined, Nakane et al. [129] also described using the test to determine the susceptibility to cracking of the consumable electrodes used in the welding process. The Murex test suffers in adoption in part due to the requirement for a large sample which is ultimately consumed during testing, as well as a requirement for a specialised test rig that allows for rotation of the specimen during welding. This has resulted in the test falling out of favour over the years for newer tests. 32 Figure 14. Murex Hot Cracking Test Schematic. Image reproduced with permission from [131]. 2.2.3.2.6 Hot Deformation Rate Test (HDR) The HDR (Figure 15) test is another test in which the weld is subjected to a tensile stress due to bending of the specimen [109]. In the test, two specimens are joined by a weld, and both are displaced during welding to create a transverse tensile deformation along the weld. A solidification crack is initiated at the start of welding, and the tensile deformation rate is reduced as the weld length increases. The point at which crack propagation stops can then be used to determine the critical deformation rate. This test is useful for the analysis of filler materials and welding parameter optimization, as the effects of the base material and weld geometry can be analysed because the specifications of the weld size can be adjusted to suit the desired dimensions [115]. The HDR test uses highly controllable and repeatable tensile deformation resulting in quantitative data in the form of a critical crack length, from which the critical deformation rate at which the crack stops propagating can be determined. Moreover, the data from this test has been found to be highly repeatable. Consistent results with clear outcomes are key goals when designing a benchmark specimen for cracking in AM. 33 Figure 15.The HDR Test. Image reproduced with permission from [16]. 2.2.3.3 Welding test summary Welding crack tests vary in size and shape, loading mechanisms, material applicability, and weld processes, and each test has inherent benefits and limitations. A summary of the welding crack tests that have been discussed is presented in Table 3. Externally loaded tests provide more in-depth and quantitative data compared to the more qualitative data obtained from self-restraint tests. However, the widespread use of self-restraint tests shows there is value in this qualitative data, as well as ways of standardising or quantifying the results of this data. This is a very important concept with regards to developing an AM hot cracking test, as external loads would be difficult to implement in some AM processes (e.g., L-PBF). When designing a self-restraint test for AM, looking at existing weld tests serves as a good first step as some geometrical features and specimens that have already been proven to be effective can be implemented, such as the cuts in the Houldcroft test or the use of multiple welds as seen in the double-fillet weld. As discussed by Goodwin [132] tests such as the Murex and Varestraint tests require a relatively large amount of material that is consumed by the test, compared to a test such as the Sigmajig test which uses much smaller coupons. Another factor in the applicability of a test is the varying complexity of geometries, such as the complex geometry Houldcroft test which adds to machining costs. 34 Table 3. Summary of described welding tests Test Applied Force Loading Loading direction relative to weld Test Output Houldcroft Self-restraint - - Crack length Fan-shaped Self-restraint - - Crack length Double Fillet Weld Self-restraint - - Crack or No Crack CPPT Self-restraint - - Circumferential Cracking (%) Sigmajig External Load Tensile Transverse Fraction of specimen length cracked vs applied stress PVR External Load Tensile Longitudinal Critical tensile deformation speed CTW External Load Tensile Longitudinal or Transverse Critical strain and critical strain rate Varestraint External Load Bending Longitudinal or Transverse Critical strain and critical strain rate Murex External Load Bending Longitudinal Bending rate vs crack length HDR External Load Bending Transverse Critical deformation rate 2.2.4 Existing solidification crack research in additive manufacturing Solidification cracking has been well documented in nearly all fusion-based AM processes for several important engineering alloys, such as 2XXX series Al alloys [133]–[135], Al6061 [65], [66], [136], [137], Al7075 [138]–[140], Hastelloy X [74], [138], [141]–[143], Rene 104 [144], and Inconel alloys [48], [79], [145]–[147]. Despite the extensive background on solidification cracking in AM, it is difficult to compare results across studies. For example, both Mazur et al. [79]. and Wang et al. [147] assess the effects of process parameters on formation of microcracks in Inconel alloys fabricated using L-PBF. On one hand, Mazur et al. found that process parameters did not substantially affect formation of microcracks in 10 × 10 × 4mm3 prismatic Inconel 625 samples. On the other hand, Wang et al. found that simulation and experimental results indicated an optimal zone of process parameters in which microcrack formation was greatly reduced for 10 × 3 × 10 mm3 samples. In both studies, crack severity was calculated as the ratio of the total area of cracks to the total sample surface area in a plane perpendicular to the build direction, yet the findings are contradictory. This could be due to the scan strategies used in each study, slight difference in alloy composition or powder size, or several other factors. In such a case, it is difficult to compare the studies as they were not produced using the same conditions. If instead, the studies used the same geometric sample with the same scan strategy, a better conclusion on the effect of process parameters on crack 35 formation could be drawn with a more rigorous analysis. Kouraytem et al. [66] studied the effects of process parameters on solidification crack propagation in AA6061 and measured the crack propagation with crack dimensions such as length, width, and depth. Many studies quantify cracking with crack density [48], [65], [79], [137], [138], [144], [147] or crack length [66], [148], [149], and others quantify cracking through qualitative means (cracking/no cracking, reduction in cracking, broad number of cracks, etc.) [139], [140], [146]. Considering the variety in quantifying results, it becomes difficult to compare results directly, as seen in Figure 16. Figure 16. Sample of different crack quantification methods. Kouraytem et al. [66] and Nie et al. [150] both quantified crack susceptibility through experimental analysis as seen in (a) and (c) respectively, whereas Tomus et al. [141] used theoretical modelling to predict hot tearing sensitivity, as seen in (b). Although both (a) and (c) used quantified experimental results, the crack dimensions used are different and thus the findings are difficult to compare. Images reproduced with their respective permissions. Multiple crack reduction strategies are also studied as methods of improving printability of alloys, however without consistent methodology to evaluate cracks, the effectiveness of each strategy cannot be easily compared. For example, Mehta et al. [65] and Uddin et al. [137] both assessed methods of crack mitigation for Al6061 produced using L-PBF, Mehta et al. looking at alloy composition adjustment and Uddin et al. looking at bed pre-heating, and both quantified cracking based on crack surface area relative to sample surface area perpendicular to the build direction. Both studies used cubic samples, Mehta et al. using cubes with a side length of 12 mm and Uddin et al. using cubes with a side length of 10 mm, and although these samples are fairly similar in dimension, the samples used by Mehta et al. have a roughly 53% greater volume and this difference affects thermomechanical behaviours such as thermal gradients, thermal cycling, and cooling times, which in turn can affect the development of internal stresses and crack formation. These 36 cuboid samples are also very simple geometrically. A benefit of AM is the ability to create complex geometrical structures, and if the tested specimen does not contain the same complex geometries that are expected in a practical part, then the reliability of the results may be limited. Crack mitigation methods might work for simple prismatic shapes, but complex shapes including high-angle corners might not see the same results. Another process parameter that lacks standardisation is the scan strategy. Mehta et al. employed a stripe pattern for the scan strategy, whereas the scan strategy for Uddin et al. is unknown. Prasad et al. [143] showed that for Hastelloy X, scan strategy variation can have an impact on crack density, but without a consistent strategy across studies the effects of individual parameters are difficult to compare as it might be a variation in another parameter that is affecting the outcome. If, instead, scan strategy is the parameter of interest, then the process parameters such as scan speed and energy density, as well as the sample geometry, need to be consistent. Other factors that affect crack formation such as minor alloying elements are also dependent on process parameters as demonstrated by Tomus et al. [141] in SLM-produced Hastelloy X. Similarly, Nie et al. [151] investigated the effect of Zr content on the microstructure of SLM-produced Al-Cu-Mg-Mn alloys and found that Zr content has a varying effect on crack density based on the scan speed and the optimal scan speed for each Zr concentration is different. Crack mitigation methods such as pre-heating [137], [146], grain refinement [65], [77], and scan strategy [79], [143], are all affected by process parameters and if these parameters are not standardised, the effectiveness of comparisons between crack mitigation methods can be reduced. Crack quantification and process parameters are not the only area where a lack of standardisation impacts reliability, as the methodology in analysing samples also varies. Locations of crack analysis and number of locations assessed both vary between existing studies and complicates the ability to draw any comparisons. As seen in Figure 17, each study is able to show the effects of their specific parameter adjustments, however when comparing the results across studies it can be difficult to interpret which crack mitigation method is more successful, as the methodology of crack analysis between studies varies. Another complicating factor when evaluating crack susceptibility is that the cracks are very small, usually with a crack opening width on the order of a few microns, which is in contrast to cracks observed in casting and welding tests which can often be identified and measured by visual inspection. Consequently, many AM studies use optical microscopy or even SEM to locate and measure and assess cracking. For example, combined SEM and EBSD is often required to evaluate the crack surface (i.e., dendritic or faceted) and location with respect to the grain structure, in order to confirm and differentiate solidification cracking from other forms of cracking (e.g., liquation cracking and ductility dip cracking). A further complication for crack assessment in AM is that unlike casting and welding, where there is generally a single crack, AM can often have several cracks in the field of view (Figure 17) since the printed material is made up of several overlapping melt tracks. For 37 tests that assess cracking based on the area density of cracking, an issue arises concerning the size of individual cracks. An area with few, but large cracks, would have the same crack density as an area with many smaller cracks. When quantifying cracks based on length, a question is raised regarding what is more important - total crack length for a given area or maximum crack length. This begs the question as to whether or not these circumstances can be considered equal when recording the data. For tests that assess based on crack dimensions, when there are many cracks, it becomes less efficient to record the data for each crack which could have a very broad range of data points and in such a situation recording results based on area would be more efficient. These issues further emphasise the need for a standardised metric of crack analysis that accounts for multiple variables such as the number of cracks, the severity of the cracks, and the susceptibility of the region to cracking, such as the crack susceptibility index employed in the CRC test. A standardised index for crack quantification could then be adopted for consistent results in an AM crack test. Figure 17. Sample of crack analysis images from four studies. Each study uses a different scale bar at different magnifications. Cracks seen in (c) produced by Tomus et al. [141] may be too small to be seen in images (b) produced by Mazur et al. [79]. Each image covers a different total surface area, and so images seen in (a), produced by Wang et al. [147], and (d), produced by Zhang et al. [133], may have a greater sample size of the total surface area compared to images in (c), however the resolution is not as high and the accuracy of crack area may not be as consistent. Images reproduced with permission. These inconsistencies further show the need for standardised methods of solidification crack analysis so as to maintain consistent samples, parameters, assessment methods, and result quantification. If these aspects of testing were standardised, the study and comparison of solidification crack susceptibility and mitigation 38 methods would be improved to a level closer to what is seen in traditional manufacturing industries such as casting and welding. 2.2.5 Existing additive manufacturing benchmark specimens Several benchmark specimen designs currently exist for AM processes. The National Institute of Standards and Technology (NIST) AM BENCH initiative outlines three main categories of benchmark specimens [152]: those developed to generate rigorous data to support the development of AM simulations and models; those to evaluate geometric features of the printer (e.g., minimum overhang angle); and those used to assess the functional performance of parts produced by AM. Three categories of AM benchmark specimens were also put forward by Mahesh [153]: geometric, mechanical, and process. Geometric benchmark specimens make up the bulk of existing benchmark tests, but there are a number of tests existing for the other two categories. Jayaram et al. [154] proposed specific mechanical properties to be tested, including the tensile and compressive strengths, shrinkage, curling, and creep characteristics; Figure 18 outlines examples of benchmark components that assess these properties. Mahesh describes process benchmarking as not only the comparison between methods, but also the analysis of the strengths and weaknesses of each and the methods used to find a balanced approach to finding the desired results [152]. Figure 18. Mechanical benchmarks as visualised by Mahesh. Image reproduced with permission from [153]. Rebaioli and Fassi [70] conducted a review of geometric benchmark specimens for a number of AM processes, which highlighted the great variance in design approaches to the problem. Benchmark designs can range from the relatively simple shapes seen in the design proposed by Iuliano et al. [155] (Figure 19a) to more complex shapes such as the design proposed by Lart [156] (Figure 19b). 39 Figure 19. (A) Benchmark proposed by Iuliano et al and (B) benchmark proposed by Lart. Image (a) reproduced with permission from Iuliano et al. [155]. Image (b) reproduced with permission from Rebaioli and Fassi [70] The need for standardisation was addressed early on by Richter and Jacobs [157] who proposed a set of guidelines for geometric benchmark specimens to set an “ideal accuracy test part” which included: (i) incorporation of several geometric features with varying dimensions; (ii) requires a relatively short print time; (iii) requires a relatively small volume of material; (iv) contains geometric features that are easily measurable; (v) includes geometric features likely to appear in a real part; and (vi) is large enough to test the capabilities at the perimeter of the build plate, not just the centre. These guidelines proposed by Richter and Jacobs [157] were intended for geometrical benchmarks, but many are applicable to other categories of benchmark specimens. For example, reduction of print time and material usage is a common goal in AM processes [158]. The idea of a specimen being easy to measure can be seen in the design of the aforementioned hot crack tests such as the Houldcroft, Varestraint, and double fillet weld tests where the location of the crack is very clear, and the desired data (crack length) is easy to evaluate. The guidelines related to geometry such as the desire for specific shapes are less applicable to evaluating solidification cracking, but knowing the geometric constraints of a machine is important for designing the specimen geometry to print. Designing for real world applications is also important, as while a specimen may not be susceptible to cracking with a simple geometry, the loads introduced with more complex geometric features could lead to cracking and, therefore, a real part could be printed without knowing the true risk of cracking. Therefore, a wide array of dimensions is desired, similar to a geometrical benchmark. 2.3 Discussion 2.3.1 Test criteria There are many factors that should be accounted for when designing a solidification crack susceptibility test or benchmark specimen for AM, but not every factor can be prioritised simultaneously. As a starting point, a crack susceptibility test will likely be a self-restraint test since these tests tend to be simpler and, (a) (b) 40 similar to casting, applying an external load to the liquid or semi-solid metal may be difficult, especially if the AM process is occurring within a small and controlled build volume. In this section, we propose a prioritised list of criteria for designing a solidification crack susceptibility test for AM. This list takes into consideration the unique challenges facing AM benchmarks and the goals of the specimen. This list is built upon the current needs of AM benchmarking and standardisation and focuses largely on the widespread adoption of a robust design. The criteria are as follows: (i) standardisation; (ii) simple and robust design; (iii) ability to assess effects of process parameters; (iv) ability to assess effects of alloy composition; (v) applicability to multiple vendors; (vi) applicability to multiple AM processes. Standardisation: Standardisation is the main objective of a solidification cracking test for AM, as analysis of cracks within this field has been conducted with a wide variety of test parameters, sample dimensions, number of sample locations, and using different measurements of cracking (when cracking is quantified), making it difficult to compare results between studies. Standardisation should focus not only on the benchmark specimen design itself, but also the methods in which cracking is measured and analysed. In addition to standardising crack detection, the quantification of cracking (e.g., crack/length or crack/area) needs to be consistent and well-defined. The area in which these cracks are assessed within the specimen, as well as how many areas are inspected, should be standardised so that geometric effects on cracking are consistent between tests. Simple yet robust design: A good design will naturally be as simple as possible, but the importance of simplicity is linked to many other factors. A simple design will likely produce results which are simpler and more easily compared among experiments, aiding in standardisation. A simple benchmark design also increases the potential range of AM processes to which it can be applied as process-specific complexities and limitations can be mitigated through simple design. For example, a test or specimen design that contains fine or geometrically complex features may be suitable for L-PBF but may then reduce its applicability to WAAM. A complex design may have features that confound the material crack susceptibility and obfuscate solidification cracking with other print defects. However, a key benefit of AM as a process is its ability to print complex geometries. The design must be robust so as to capture complexity that would be expected from real-world parts. It is important to keep in mind that the test serves as a tool to be applied to real manufacturing processes, and an overly simple design might lack critical details that could affect the real-world counterpart. Simplicity would also mean a reduced risk of failure during printing, reducing test time and costs, and also increasing the likelihood of widespread adoption. Material costs for AM are still quite high, particularly for the fabrication of custom composition powders or wires, and any economic optimizations are valued. Consequently, the amount of material consumed should be minimised. 41 Assessing multiple parameters: A successful test must be able to assess different materials under the same process parameters, since alloy composition is strongly linked to crack susceptibility. This means that the test would need to differentiate the printability of different material classes, as well as be sensitive enough to reliably detect any variance in crack susceptibility between similar alloys. The ability to evaluate compositional effects on cracking is particularly important for AM, as there is a currently growing interest to design new alloys specifically for AM with previously underexplored alloying combinations, such as aluminium-rare earth systems. Assessing alloy composition: Process parameters are not the only input that affects crack susceptibility. The material used is an important distinction and a test would ideally be able to assess different materials under the same process parameters. This means the test would need to consider printability of various materials, as well as be sensitive enough to reliably detect any variance in crack susceptibility between similar alloys. Applicability for multiple vendors: A benchmark design should give reasonably consistent crack results for the same process parameter inputs and material feedstock regardless of the printer used to print it for a given category of AM processes (e.g., L-PBF or L-DED). In other words, the results of the benchmark specimen design should transcend AM system vendors. Applicability for multiple AM processes: Ideally, an AM crack test or specimen would transcend the specific AM process and heat source. A test that is designed for L-DED could likely be used for electron beam-DED without much, if any, modification. However, if this test was able to be printed using other types of AM processes such WAAM or L-PBF then this opens up a new layer of possible analysis. Not only does this increase the likelihood of widespread adoption for a specimen, but it also allows for comparisons between not only materials and print parameters, but also between entirely different processes. This can allow for the evaluation of processes to determine which would be the most reliable for a certain application or product. This is analogous to existing tests such as CTW and PVR that are capable of quantifying cracking in both arc and laser welding. However, it may not be possible or even necessary for initial test designs to be applicable across all types of AM processes. Even within the more mature fields of casting and welding no test exists that fully characterizes cracking for all processes, and modifications often need to be made, such as the development of the fan-shaped test through adaptation of the Houldcroft test. This is not an exhaustive list of what is necessary for any given test. Rather, it provides a broad framework to structure design choices when developing a solidification cracking test or benchmark specimen for AM, and these concepts can easily be adapted to other types of benchmarks such as geometric or mechanical. Table 4 outlines several additional factors for consideration that increase complexity in test development 42 for AM. A number of specific concepts arise when designing for AM, and though a test design might not attempt to address each of these concerns, they are still areas of consideration when deciding on the objectives of a test. Examples of existing casting and welding tests that address these concepts are included in the table. Table 4. Additional considerations when designing a solidification crack test for AM. Concepts for AM test Existing Casting or Welding Tests Build orientation PVR, Varestraint/Transvarestraint Heat transfer through the specimen CRC, Critical Diameter Internal geometric structures Unique to AM Quantification of qualitative results CRC, CPPT Standardisation of assessment ECRC, Varestraint, Sigmajig, Double fillet Instrumentation (sensors, thermocouples, load cells, etc.) CAST test, ECRC, externally loaded tests Variation in loading Unique to AM 2.3.2 Assessment of existing tests Ideally, an existing casting or welding crack susceptibility test could be used or adapted for use in AM, as these tests are already familiar to those concerned with solidification cracking and there is historical crack data available for comparison to determine if measurements in AM samples are reasonable. AM is limited in a similar way to casting by way of relying largely on self-restraint tests, due to the difficulty of applying a controlled load to a liquid or semi-solid material, especially for AM where the melt pool is often very small and continually changes locations and directions throughout the build. With respect to the additional considerations in Table 4. that might need to be addressed in AM and comparing them to the concepts found in existing tests, design ideas that have already proven to be effective can be drawn from welding and casting and adapted to AM. In this section, we will assess the existing casting and welding tests highlighted in this review against the proposed criteria for an AM solidification benchmark design outlined in section 4.1. The three main tests considered here are the CRC, CDM, and the Houldcroft tests, as they are all self-restraint tests, which is most suitable for a novel AM test, and they all employ design ideas that provide valuable guidance in standardisation and the promotion of cracking using geometry alone. As noted above, the standardisation and quantification of crack assessment is of highest priority for a benchmark specimen. Nearly all of the casting and welding tests that were previously described provide a standard or systematic method to assess crack susceptibility and many of these tests provide some quantitative method to unambiguously assess cracking within the constraints of the test. By example, the 43 Houldcroft, fan-shaped, CPPT, constrained ring, and Warrington and McCartney test all evaluate crack susceptibility by the maximum crack length when no external load is applied during solidification. Other tests further increase standardisation by specifying predetermined locations to assess cracking, such as the CRC test, and by following standardised process procedures and ensuring satisfactory completion of the procedure, as seen in the double-fillet test. The CRC test in particular has an excellent level of standardised procedures for evaluating cracking. Pre-defined crack severity ratings and locations of cracking, as well as a formula to quantify the overall crack susceptibility, is an excellent method of creating repeatable results across studies. The test design is also relatively simple and can be used to assess various parameters. However, this design is not directly suitable for AM as the design relies on the mould geometry to impose reactive forces against the contracting metal. In this sense, the geometric design of the test cannot be translated to AM directly, but the concept of a formula to standardise results and defined locations of cracking could be instrumental in the formulation of a benchmark specimen or test. Similar to the CRC test, the CDM test uses the idea of evaluating crack susceptibility through a common geometric feature (rods in the case of both CRC and CDM) in which the geometry is repeated across the specimen with one dimension varying to change the level of restraint on the solidifying metal (diameter and length for the CDM and CRC, respectively). Again, although the CDM design cannot be used in AM directly, the concept of controlling restraint on the liquid metal by changing the dimension of a simple geometric feature is a potential design idea which could be the basis of an AM specific test. The Houldcroft test, which uses the geometry of the substrate to affect its stiffness and, thus, its crack susceptibility, is another case of a test geometry not being directly transferable to AM without modification, but the general ideas behind it being very applicable. This test is standardised, simple to perform and interpret (i.e., maximum crack length), assesses multiple parameters and is applicable for various welding processes. However, the design itself is specific to welding. If a new design was developed for AM which took into consideration these features of the Houldcroft test, it could be as widely used within the AM industry as the Houldcroft test is within the welding industry. None of the existing tests can replicate the fundamental layer-by-layer aspect of AM processes and, as such, cannot fully account for the complexities inherent to AM. However, the ideas from these three tests (CRC, CDM, Houldcroft) can form a design theory that can be used to dictate the process of designing an AM cracking test. Creating a test that has a defined and standardised measurement for crack susceptibility which includes a formulised index to act as a comparative result among tests will help achieve the first criteria of standardisation. The simple designs used by these three tests serve as a guide on how to simplify testing for 44 AM; whether using ‘levels’ of a geometric feature to assess susceptibility, or a single specimen with varying susceptibility across its length, these ideas can be adapted into a test for AM in order to create a test sensitive to multiple parameters while remaining simple enough to understand and apply to various AM systems and processes. Each of these tests is a self-restraint test which indicates the value of a self-restraint test is not overshadowed by the additional numerical data obtained from an externally loaded test. Especially as a novel test, a self-restraint test can fill the broader gaps in AM solidification crack benchmarking without overcomplicating the design to suit more specialised goals. These core concepts can be used as an initial guideline when developing an AM hot crack benchmark to aid in the success of its design. 2.4 Conclusions and future outlook Solidification cracking, or hot tearing, is a metallurgical defect occurring toward the end of solidification and causes a serious impairment in mechanical properties for cast, welded, and 3D printed metals. The fundamental mechanisms of solidification cracking remain consistent between casting, welding, and additive manufacturing. However, there are unique process conditions that dictate viability of solidification crack susceptibility testing methods. Each manufacturing process has certain characteristic features and limitations that restrict the testing processes that can be applied, such as the difficulty of applying external loading during casting or AM. A benchmark specimen allows for a reliable way to quantify crack susceptibility for a given process in a way that accounts for the process’s benefits and limitations. Without a consistent method of testing, process and material development is hindered and makes quality assurance much more difficult and therefore slows the industry adoption of AM. The following conclusions can be drawn from a review of the existing literature: 1. No existing solidification crack test can be readily adopted for AM. However, certain design concepts, such as a repeating geometry that varies in a critical dimension, can provide useful starting points. 2. Existing methods of crack susceptibility quantification, such as susceptibility indexes and standardised crack inspection methods, are viable candidates for transferability to a unique AM test. 3. Self-restraint tests are the best candidate for an AM crack test design, as there are many challenges involved with implementing an external load and self-restraint tests have been shown to provide valuable results despite relying entirely on the inherent stress risers induced by the specimen geometry. Determining what factors of solidification crack susceptibility are to be tested is the first step in designing a test specimen, as the rest of the test will rely on what data needs to be obtained. AM has a number of 45 factors unique to its processes, and as such existing casting and welding tests cannot be transferred directly to these processes. For an initial AM crack test, a broad, general approach to testing crack susceptibility is valuable as when no other test exists, a starting point for more specialised tests to build off can allow for further development in the field. The focus of a general test should be on the standardisation of crack analysis and quantification in order to bring the level of standardisation closer to that seen in welding and casting tests. These concepts of standardisation can be translated from existing tests to AM, even if the test designs themselves are different. The evolution of solidification crack tests in casting and welding sheds light on the factors to consider when looking for a standardised method for widespread adoption. Key factors include clear definitions of crack severity and how to quantify it, as well as pre-defined inspection locations, all to help standardise results between studies. More complex factors in crack testing can be the subject of future iterations, and building a steady foundation for future tests is a valuable endeavour in the current state of crack testing. The outcomes of this review highlight important features for standardisation and test design in AM for not only crack tests, but benchmark tests as a whole. The need for standardisation and easily understood test methods is a clear barrier that needs to be addressed in AM research in order for it to be elevated to the same standards and industry acceptance as traditional manufacturing methods. In order to highlight the advantages of AM, the results of research must be verifiable and consistent so that a high level of confidence can be maintained for industry development. The difficulties in developing a test must be addressed by future research and creates a wide field of study focussed on addressing each of these challenges. Such challenges include the wide range of AM processes which in turn have their own unique set of challenges and process characteristics. In a general sense, the complex thermomechanical history induced by AM processes and the layer-by-layer method is a challenge not faced by traditional manufacturing methods. A robust design will need a combination of experimental data and numerical simulations in order to address such a complex challenge. Depending on the process, cracks may be very small and require advanced imaging techniques, such as scanning electron microscopy (SEM) as opposed to optical microscopy, which may hinder practical use of a given test and limit its adoptability. Having pre-defined locations for crack inspection can help to counteract this, and future designs should account for this potential method of standardisation. Other challenges such as accounting for the complex geometries that are possible with AM processes, and the limitation of external loads to induce stress in a benchmark, are areas of interest that will need to be addressed as development of AM benchmarks progresses. The development of these tests is a critical step towards integrating AM as a regular part of the production process and increasing industry adoption. 46 3.0 Objectives The research presented in this thesis aims to address the critical knowledge gap in that no standard, quantifiable method of SC susceptibility exists for AM processes. Although research into the mechanisms of SC and methods of SC reduction in AM is ongoing, methods of result quantification and analysis vary greatly across studies. This greatly hinders the ability to compare results between studies and slows down the adoption of AM processes in industrial applications. Thus, the objectives of the research in this thesis are as follows: i) Use the assessment of the current state of solidification crack tests in both welding and casting in order to determine features best suited for an AM benchmark and guide initial specimen design. ii) Design a novel benchmark specimen that can produce quantifiable results on solidification crack susceptibility for AM processes. iii) Develop a solidification crack susceptibility index to accompany the benchmark specimen developed in objective (ii) in order to further improve the quantification of results. This research aims to provide a foundation for further development of SC susceptibility tests for AM processes, as the complexity of the many AM methods are unlikely to be fully encapsulated in a single test. However, the benchmark specimen that will be developed aims to provide a simple and robust test that can be applied to multiple materials and processes, as opposed to focusing on any specific AM process. This provides a direction for further research and development in this topic while addressing key issues in the current state of research such as the lack of standardization in SC studies. This thesis will address these knowledge gaps, provide an initial step in addressing them, and discuss the feasibility and constraints of the proposed benchmark specimen. 47 4.0 Initial Crack Benchmark Specimen2 4.1 Introduction The objective of this chapter is to develop a novel SC susceptibility benchmark designed for AM processes based on the knowledge gap identified in Chapter 2. As this is a novel test that does not build off an existing AM design, this benchmark serves as a first step in the development of future tests. Consequently, the benchmark design developed in the current study was intended to be a simple yet robust design so as to be applicable to a wider range of processes and conditions to provide a basic level of repeatability among studies and to provide insights into the steps required to improve future designs and increase standardization. The benchmark design was printed by L-DED and the development of this benchmark sample required several printing trials, exhaustive microstructure characterization, and finite element (FE) simulations of the L-DED printing process. 4.2 Materials and methods 4.2.1 Benchmark design In order to maintain a simple but robust benchmark, the initial design was intended to use geometric features that could be produced using a range of AM processes and, as such, aimed to remove features that require supports (e.g., large overhangs, internal cavities) or high geometric precision (e.g., thin sections). The geometry for the resultant design that was used in the current study is shown in Figure 20. The basic design philosophy was to print a number of individual ‘legs’ that act as a restraint on the metal undergoing solidification in layers above the apex of where the legs are joined. This geometrically induced self-restraint is intended to impart a tensile load to the melt pool to promote solidification cracking. Results presented at the recent National Institute of Standards and Technology (NSIT) AM Bench 2022 conference on the AMB2018-01 residual stress benchmark challenge, which featured a bridge-like design similar to the crack benchmark presented in Figure 20, typically showed high levels of longitudinal strain along the edges of the top surface [159], [160]. Therefore, the current design employs a tapered peak at the top of the design which was intended to vary residual stress based on location along the sample and to promote SC at the apex of each leg intersection. Utilizing these findings, assessment of crack severity would then ideally account for both the crack length and location (i.e., apex). Cracks observed further away from the high-stress regions would carry a higher weighting of severity compared to cracks of the same length found closer to high-stress locations. Initial estimations suggested the central apex to be the location of highest 2 The contents of this chapter have been adapted from the original manuscript prepared for submission: A. Wall, T. Dong, M.J. Benoit, “On the Development of a Novel Benchmark Design for Crack Quantification in Additive Manufacturing”, Additive Manufacturing (submitted July 2023). 48 stress concentration as it has the most constraint, imposed by the rest of the specimen, compared to the other apexes. Verification of this hypothesis would be based off simulation of the process, as well as initial findings. The proposed design and quantification method were inspired by several existing tests and benchmarks, both for AM and conventional manufacturing processes. The idea of using a repeating geometry to induce varying levels of stress at different locations was adopted from existing self-restraint tests in casting and welding such as the varying rod length in the Constrained Rod Casting (CRC) test [27], and the varying slit length in the Houldcroft Test [25], respectively. These existing tests provide insight into the utility of self-restraint tests, as well as methods of quantifying the results such as the crack susceptibility index used in CRC which assess hot crack severity based on not only the severity of the crack, but also which rod the crack occurred in, as longer rods are more susceptible to cracking [27]. Figure 20. Design schematic for the benchmark specimen with a 10mm thickness. All units are in mm. 4.2.2 Materials Gas atomized Inconel 625 (IN625) powder supplied by Kennametal with a nominal particle size distribution of 45 – 90 µm was used in the current investigation to print the majority of the benchmark samples. Inconel 625 was used as it has been shown to experience SC during AM [79]. Additional samples were also printed with gas atomized Inconel 718 (IN718) powder, also supplied by Kennametal with a nominal particle size distribution of 45 – 150 µm, in order to assess the ability of the benchmark to evaluate crack susceptibility between similar alloys, as IN718 is known to be less crack susceptible than IN625 [161]. The chemical composition of both powders as measured by inductively coupled plasma optical emission spectrometry (ICP-OES) is presented in Table 5. Scanning electron microscope (SEM) images of the as-received powder and powder cross sections are shown in Figure 21. From the figure it can be seen that the as-received powders are fairly uniform in shape, with mostly spherical particles, though some variation is observed with the inclusion of ellipsoidal particles. The particles also have several small satellites adhered to them, with varying number and size. The cross-section of the powder particles revealed a dendritic structure with 49 interdendritic phases enriched in Nb and Mo, and smaller, dark spots rich in Ti. Pores were observed in some particle cross-sections, though they were relatively uncommon. Table 5. Chemical composition of powders as measured by ICP-OES (wt%) Ni Cr Mo Fe Nb+Ta Co Ti Al Mn Inconel 625 bal. 21.23 8.71 0.29 2.38 0.08 0.04 0.15 0.29 Inconel 718 bal. 19.65 3.16 18.50 4.83 0.03 1.13 0.57 0.20 Figure 21. SEM images of the (a, c) as-received powder morphology and (b,d) the particle cross-sections for (a, b) IN625 and (c, d) IN718 powders. The small dark spots in (b) and (d) were determined to be phases rich in titanium. 50 4.2.3 Benchmark fabrication The benchmark samples were manufactured using laser directed energy deposition (L-DED) on a Liburdi Automated Welding System (LAWS) 1000 with a spot size of 0.9 mm, supplied by Liburdi Automation Inc. A bi-linear stripe scan strategy was used, with the scan vectors alternating in direction by 90° every two layers, as demonstrated in Figure 22. The laser power varied with layer number based on vendor recommendation ( Figure 22. Scan strategy of the printed samples where n = the first layer. The scanning direction alternates by 90° every two layers. Table 6), as the samples were printed on a large heat sink (substrate) and so the first (root) layers were printed at a higher power to ensure complete fusion, with transitional layers to gradually lower the power (transition layers) before using a lower power in the remaining (fill) layers. All other process parameters were held constant based on vendor recommendation and are presented in 51 Table 7. All samples were printed on 304 stainless steel substrates with dimensions 203 × 203 × 6.35 mm3. Two independent replicate benchmark samples were printed from IN625 for repeatability analysis, and a single benchmark was printed from IN718 for comparative purposes. Figure 22. Scan strategy of the printed samples where n = the first layer. The scanning direction alternates by 90° every two layers. Table 6. Laser power variation between layers Layer Power (W) Root (Layers 1 – 2) 350 Transition (Layers 3 – 4) 325 Fill (Layers 5+) 300 52 Table 7. L-DED process parameters Process parameter (units) Value Powder Feed Rate (g/min) 2.5 Travel Speed (mm/min) 508 Layer Height (mm) 0.4 Ar shielding gas flow rate (L/min) 11.8 Ar nozzle gas flow rate (L/min) 2.8 Ar powder carrier gas flow rate (L/min) 3.5 4.2.4 Microstructural characterization and analysis The benchmark specimens were cross sectioned lengthwise at the mid-plane of the sample after printing in order to assess cracking at the apex of the legs. Cross sections were mounted in conductive resin and prepared by standard metallographic grinding and polishing, using a final polishing step with 0.04 µm colloidal silica. The samples were not etched after polishing, but the colloidal silica produced a slight etching effect. A Keyence VHX-970F was used for optical microscopy. A Tescan Mira 3 XMU SEM equipped with an Oxford Instruments X-Max energy dispersive X-ray spectroscopy (EDS) detector and an Oxford Instruments Nordlys electron back scatter diffraction (EBSD) detector were used for detailed microstructure analysis. SEM imaging, in both secondary (SE) and back scatter (BSE) electron modes, and EDS measurements were performed with an accelerating voltage of 15 kV, while EBSD measurements used an accelerating voltage and step size of 30 kV and 1µm, respectively. X-ray diffraction (XRD) analysis was performed on the Inconel 625 benchmark samples to determine possible phases using a Bruker D8 Advance system. All measurements were taken using a Cu anode (K𝛼1 = 1.5406 Å) with a Ni K𝛽 filter. Samples were randomly oscillated in the X/Y plane to increase particle statistics. The sample was rotated through a 2θ range of 20 º to 100 º with a step size of 0.03° and step time of 2s. Phase identification was performed with Difrac EVA (version 6.1.0.4) using the Crystallography Open Database (rev.269086). 4.2.5 Simulations 4.2.5.1 Solidification simulation The solidification behaviour of both IN625 and IN718 was simulated using ThermoCalc (version 2022.b) with the TCNI12 Ni database. The non-equilibrium (i.e., Scheil) module was used for the simulations to predict the temperature vs. solid fraction (T vs. fs) curves and phase evolution during solidification. A temperature step size of 1 ºC was used, with no fast-diffusing elements or back diffusion selected. No phases 53 were suppressed during the simulations. The material compositions used for this analysis is based on the ICP-OES results outlined in Table 5. 4.2.5.2 L-DED process simulation The directed energy deposition module of the commercially available Simufact Welding (Hexagon MSC Software, version 2022.0.1) finite element (FE) software was used to simulate the thermomechanical behaviour of the benchmark specimen throughout printing and to provide a prediction of the temperature and stress history at crack locations. The FE mesh for the benchmark was comprised of hexahedral elements with a size of 0.4 mm, resulting in one element layer per printed layer, and was produced with the software MSC Apex (version 2023.1). The mesh for the substrate was similar, but used an element size of 1 mm. Nodes at each corner of the substrate that were fixed in the x, y, and z directions served as the boundary conditions with the first layer of the printed specimen being fixed to the substrate. The material for the benchmark specimen was designated as Inconel 625 and the baseplate was designated as 304 stainless steel, using the thermophysical properties in the Simufact database for each material. The toolpath used for the simulation was created using the GCode produced by the Liburdi LAWS software for printing of the benchmark geometry (section 4.2.3 Benchmark fabrication). These simulations used the exact input geometry based on the CAD files from which the mesh was built, and such they cannot account for geometric inaccuracies that may occur during printing. The simulations were performed using the ‘Advanced Thermal Cycle’ setting in Simufact in order to reduce calculation time due to the size and complexity of the simulation. This setting combines heat input for a defined number of timesteps, followed by a cooling period. For the current simulation, 25 timesteps were merged for a given calculation, which resulted in a calculation approx. every 2 seconds of simulated process time, and a single calculation could not contain more than one weld track. This setting may negatively impact the quality of results, especially due to the complexity of the geometry; thus, the absolute values of thermomechanical properties produced by the simulations may have an unknown margin of error. Nevertheless, the Advanced Thermal Cycle setting was required for the simulations to reduce computational time (~1 week) and file size (~1 TB), and the level of error associated with using this simplification will be discussed later. Consequently, the FE simulation results were used to generally understand the local thermomechanical history throughout the benchmark sample during printing rather than to predict the true values. 4.3 Results 4.3.1 Crack quantification The printed benchmark specimens can be seen in Figure 23 for both IN625 and IN718, in which the layer-wise nature of the L-DED process can be seen. Nevertheless, the bulk material in the samples was found to generally be of high density (i.e., low porosity) and of good quality, except for the regions in which cracks 54 were intended to form. The optical microscopy images shown in Figure 24 reveal the presence of cracks located at the apexes of where two legs join in the sample for both IN625 benchmark specimens. Note that replicate 2 was initially cross sectioned along apex 5 for analysis but no substantial results were obtained; consequently, no images were collected in the longitudinal section at this location. As seen in Figure 24, many of the cracks propagate in a similar fashion, in which the zig-zag path appears to follow the weld pool boundaries. The cracks were quantified by measuring both the maximum vertical length of the cracks from the apex of the intersection to the tip of the crack (apex-to-tip length), and the total crack propagation length (total crack length), the results of which can be seen in Table 8 and Table 9. Measurements were conducted on optical micrographs taken at 100× magnification. A crack was only measured if it was located at the apex of the intersection, and the distance measured includes any cracks or other voids that are connected to this initial crack. By example, while a void can be seen in the micrograph of apex 4 of replicate 1 in Figure 24, no crack length was recorded in Table 8 as no crack was observed to originate at the apex. The ratio of the apex-to-tip length and the total crack length was also determined. Interestingly, when cracks were observed at a given apex, the apex-to-tip:total crack length ratio was found to be between 0.8 to 0.9 for both replicates. It is worth noting that by inspection of 55 Table 8 and Table 9 that there is no clear relationship between any quantitative crack measurement and apex location in the sample. One challenge that was encountered when quantifying crack lengths was that the visibility of cracks was found to be strongly influenced by the observed plane within the sample; a deviation of a few microns further into the sample (e.g., from grinding and polishing) may reveal or remove cracks from view. This also impacted the observed width of the crack and, therefore, the observed surface area of the crack. For this reason, width and surface area measurements were not used for quantification. Comparison of the crack lengths for replicate 1 ( Table 8) and replicate 2 (Table 9) highlight the difficulty in obtaining highly repeatable crack measurements, even in samples processed from the same powder with the same parameters and analysed at nominally the same location. Furthermore, the observation of voids near some of the apexes confounds crack observations, as SC is closely related to the formation of shrinkage pores and lack of fusion defects can act as stress risers; nevertheless, the inclusion of pores in the crack quantification measurements was intended to simplify the quantification process and avoid further variation in measurements. 56 Figure 23. Printed samples of the benchmark specimen for both IN625 and IN718. 57 Figure 24. Optical microscopy images of cracks at the intersections of legs in both of the IN625 samples. Many of the cracks follow a similar 'zig-zag' path, and all cracks have a similar initial narrow crack geometry. Note that replicate 2 was originally cut transversely through apex 5 and, therefore, the logitudinal cross section at this location was not able to be imaged. 58 Table 8. Crack quantification at the apexes seen in Figure 24 for IN625 replicate 1. Apex No. Apex-to-Tip Length (μm) Total Crack Length (μm) Apex:Total Ratio 1 0 0 - 2 1579 1842 0.86 3 868 952 0.91 4 0 0 - 5 690 847 0.81 6 1460 1739 0.84 7 1256 1424 0.88 Table 9. Crack quantification at the apexes seen in Figure 24 for IN625 replicate 2. Apex No. Apex-to-Tip Length (μm) Total Crack Length (μm) Apex:Total Ratio 1 625 764 0.82 2 1197 1368 0.88 3 944 1127 0.84 4 838 988 0.85 5 N/A N/A N/A 6 784 880 0.89 7 0 0 - 4.3.2 Microstructure analysis Higher magnification SEM images of the cracks shown in Figure 24 in replicate 1 at apex 2 (Figure 25) and apex 6 (Figure 26) reveal a fracture surface morphology that differs along the crack length, but show strong similarities between the two cracks. As seen in Figure 25, the crack at intersection 2 differs in width and a change in fracture morphology occurs at one of the points where propagation changes direction. Figure 26 shows an SEM image of a similar instance of hairline fracture propagating from a larger crack observed at intersection 6, and it appears that this smaller crack propagates along an interdendritic pathway (Figure 26c). The fracture surface differs based on the location in the crack as well. As seen in Figure 25e, the surface along the large, uniform crack is rough and dimpled, whereas the surface shown in Figure 26e is relatively smooth with a number of ‘spikes’ across the surface, a morphology noted by Rappaz et al. [162] to be related to solidification cracking. The fracture surface of the hairline fractures is difficult to determine due to their narrow opening, however they do not show signs of a dendritic fracture surface that is common of SC [163]. 59 Figure 25. SEM-SE images of various locations along the crack in IN625 replicate 1, apex 2 as seen in Figure 24. Figure 26. SEM-SE images of various locations along the crack in IN625 replicate 1, apex 6, as seen in Figure 24. 60 The EBSD analysis of the microstructure in the proximity of the relatively wide and straight cracked regions in replicate 1 at apex 2 and apex 6 is presented in Figure 27. The inverse pole figure (IPF) shows that the cracks propagate in a transgranular mode , as seen by the same crystallographic grain orientations on either side of the crack, and the dimpled surface along the fracture (Figure 25e) as a dimpled surface is common for ductile fracture modes [164]. Moreover, Figure 27c and 8f show a high number of low-angle grain boundaries (LAGB) in green for the cracks at apexes 2 and 6, commonly associated with high strain rates and plastic deformation [165]. These transgranular fractures appears to occur close to the melt pool boundaries, as indicated by the aligned columnar grains orthogonal to the crack and the change in directionality of the grains below the fracture location in Figure 27b, which suggests solidification in different directions (i.e. towards the centre of their respective melt pools). IPF maps of the region around the tip of the hairline fractures at both apex 2 and apex 6 in IN625 replicate 1 (Figure 28) reveals that these cracks appear to continue to propagate transgranularly, which indicates that they are not likely to be SC, as SC are well known to propagate in an intergranular manner [34]. Figure 27. (a, d) SEM-SE images, (b, e) IPF-Z grain orientation maps, and (c, f) grain boundary maps for IN625 replicate 1 at (a-c) apex 2 and (d-f) apex 6. The maps show a common transgranular fracture mode that propagates close to the weld pool boundary, with a high density of LAGB adjacent to the cracks (shown in green in c and f). 61 Figure 28. (a, c) SEM-SE images and (b, d) IPF-Z maps of the region surrounding the hairline fractures at (a, b) apex 2 and (c, d) apex 6 in IN625 replicate 1. EDS mapping of the wide, transgranular fracture region of the crack at apex 2 (Figure 25e) reveals the presence of interdendritic phases enriched with Mo, Nb, and Cr oriented orthogonally to the direction of the fracture (Figure 29), further confirming the transgranular nature of the cracks. In contrast, the hairline region of these cracks at both apex 2 and 6 appears to propagate along these enriched interdendritic regions, as seen in Figure 30 for the crack located at apex 2. However, the IPF-Z map of this region (Figure 28) indicates that portions of this region may in fact propagate transgranularly as well. 62 Figure 29. (a) SEM-SE image and (b-e) corresponding EDS maps of a portion of the transgranular fracture for the crack observed in IN625 replicate 1 at apex 2 in Figure 24. Phases enriched with Mo and Nb can be seen to be aligned orthogonally to the crack direction. Figure 30. (a) SEM-SE image and (b-e) corresponding EDS maps of the tip of the hairline fracture at apex 6 in IN625 replicate 1 in Figure 24 . The results of the XRD analysis for IN625 are presented in Figure 31. In general, the phase results are consistent with what is expected for IN625, with peaks corresponding to the primary γ-Ni phase and minor peaks that may correspond to MC, M6C, or M23C6 carbides [166]. Minor peaks were observed but were unable to be classified using the available database. 63 Figure 31. XRD analysis of IN625 replicate 1 sample. The precipitates in Figure 30 may be these carbides observed in the XRD analysis, as MC, M6C, and M23C6 carbides in Inconel 625 are rich in Nb, Mo, and Cr [166], or they may correspond to the predicted Laves phase (Figure 33) which is also enriched with these elements [167]. However, the database could not provide an accurate indication of the presence of either of these phases, and judgements must be made with reference to previous studies. 4.3.3 Inconel 718 benchmark analysis The optical microscopy cross section images at the apexes for the benchmark specimen produced with IN718 can be seen in Figure 32. It can be seen from the figure that far fewer cracks were observed in the IN718 sample compared to the IN625 sample (Figure 24), which is in good agreement with the general understanding that IN718 more resistant to SC than IN625 [161]. Nevertheless, it should be noted that in instances where cracks were observed in the IN718 sample (e.g., apex 4 in Figure 32), the crack exhibits a similar zig-zag propagation path as those observed in IN625 suggesting a similar crack mechanism. However, these cracks were shorter than those observed in the IN625 sample and, as such, little information on crack morphology could be gathered. It should be noted that lack of fusion pores can be seen above the apex in several micrographs. In addition to the overhanging geometry in this region, the IN718 sample was printed using process parameters that were optimized for high density IN625. Consequently, the energy density may not have been optimal for IN718. By comparison of the IN625 and IN718 benchmark samples, it appears that the design of the benchmark specimen is able to repeatedly produce cracks with the same morphology at reliable locations (i.e., at the leg apex) even for different alloys within the same material class, and is able to demonstrate the effect of alloy composition on crack susceptibility. 64 Figure 32. Optical microscopy of cracks located in samples produced using IN718. The cracks are not as severe as those seen in IN625, but a general propagation path (zig-zag) similar to the previous fractures can be seen at apexes 1 and 4. 4.3.4 Solidification simulations Figure 33 presents the solidification path for both IN625 and IN718 in the form of the temperature vs. solid fraction (T vs. fs curve). The solidification temperature range of IN625 was predicted to be greater than that for IN718, and a broader range is generally associated with increased susceptibility to SC [118]. Several minor phases were predicted to form toward the end of solidification (e.g., Laves, σ), which correspond to the interdendritic phases observed in Figure 30. It should be noted that since no carbon was detected in the ICP-OES data in Table 5, the carbide phases identified in the XRD analysis (Figure 31) were not predicted during solidification in Figure 33. 65 Figure 33. Thermocalc predicted solidification behaviour of IN625 and IN718. 4.3.5 Thermomechanical simulations of L-DED The results of the Simufact simulation can be seen in Figure 34, in which the simulation has been frozen after printing of the layer that joins the apexes of the legs. It can clearly be seen in Figure 34 that stress (and strain) localization occurs throughout the sample thickness at each of the apexes. To further understand the local thermomechanical history at the apexes during printing, the temperature and equivalent stress history at the nodes at the longitudinal centreline in the top layer shown in Figure 34 were extracted from the FE simulation (Figure 35). For the sake of brevity and clarity, the thermomechanical histories for apex 2 and apex 6 are presented below, as these were the focus of the microstructure characterization above. 11001150120012501300135014000 0.2 0.4 0.6 0.8 1Temperature (°C)fsIN625 IN71866 Figure 34: Simulated (a) temperature, (b) equivalent stress, and (c) plastic strain at the completion of the layer that connects each apex. Higher levels of stress and strain can be seen along the length of each apex. Red arrows in (a) show the scan pattern orientation of this layer. 67 The temperature profile at each apex clearly demonstrates the repeated reheating and rapid cooling due to the deposition of adjacent tracks and subsequent layers. It should be noted that apex 6 was deposited first in the layer in which the apexes were first joined based on the scan strategy and layer sequence and, therefore, appears before apex 2 in the thermal history. Due to the tool path, as the deposition head travels from one end of the specimen to the other and then returns at the end of the subsequent layer, one side of the specimen will have one layer end and the next begin directly on top of it immediately, whereas the other end of the sample will have a longer period of cooling between layers. This affects the period between reheating cycles and the cooling rate of each apex and is heavily impact by the scan strategy. By example, from Figure 35, it can be seen that the time period between the initial deposition and subsequent re-heat cycle at apex 2 is approximate 30 sec, whereas the difference between deposition and subsequent heating at apex 6 is approx. 95 sec. It should also be noted that the peak temperature observed in the thermal history in Figure 35, which corresponds to melting during deposition at each apex, is approximately 1085 °C to 1100 °C. These values are below the Thermocalc-predicted liquidus temperature expected for the IN625 powder composition used in this study (Figure 33). It is expected that this discrepancy can be attributed to two main causes. First, it is well known that the melting range of IN625 is highly sensitive to the exact composition of the sample [168]. Thus, if the composition and corresponding melting temperature range of the powder used in the current study does not exactly match the values used in the Simufact material database, then the temperature could be underpredicted in the simulation. The Simufact material database defines the solidus temperature for IN625 as 1137 °C the liquidus temperature as 1336 °C, so the peak temperature should have been above 1336 °C in the simulations. While some instances of lack of fusion were noted in the printed samples (e.g., replicate 1 apex 1 and replicate 2 apex 7), high density and good fusion were clearly observed in the samples, indicating that melting did of course occur during printing. Secondly, it was noted in section 2.5.2 that the Advanced Thermal Cycle setting was used to reduce computational demands by merging data for several timesteps in the calculations, which is expected to reduce the precision of the simulation predictions. Consequently, the simulation data is used here to generally understand the thermomechanical history of the material with an expected level of inaccuracy. The evolution of the equivalent stress at apex 2 and apex 6 can be seen in Figure 35. As previously noted, apex 6 was printed before apex 2 and, as such, had a longer cooling period compared to apex 2 as the heat sourced moved away towards the other side of the specimen, and then returned. This resulted in a larger difference in maximum and minimum temperatures experienced at apex 6 compared to apex 2, which in turn resulted in a variance in stress experienced at each apex. Because there is a greater duration between melting & re-heat at apex 6, it can cool more and, consequently, experiences more thermal contraction. Moreover, the cooler underlying metal would have a higher strength compared to just-deposited material 68 (e.g., at apex 2), so the degree of restraint at apex 6 during the first re-heat is greater. These changes are reflected by a greater equivalent stress at apex 6 compared to apex 2 in Figure 35. This highlights the impact of scan strategy on the outcome of this specimen, as the direction of scanning can vary based on layer number as well as the location along the specimen in a given layer. It should be noted that although stress is predicted to be lower in apex 2, cracks were still observed at both locations. Figure 35. Simufact Welding simulated temperature and equivalent stress vs. print time of (a) apex 2 and (b) apex 6. The effect of the scan strategy on the temperature profile and stress evolution can be seen in the difference in reheating frequency at each apex as the layer number increases. Although stress development progresses as the remaining layers are deposited and the part begins to cool upon completion, for the purpose of a solidification crack test, the residual stresses present in the solidified material are not of significant importance, as any fracture induced in the solid-state will not be a solidification crack. Although these solid-state fractures may hinder result quantification for SC, the most important time for stress development is during the final stages of solidification as the material is deposited, as this is when residual stresses may initiate a solidification fracture. 4.4 Discussion 4.4.1 Effectiveness of benchmark design From the above analysis, the benchmark design successfully produced cracks that were easily observed with optical microscopy in specific, predetermined locations in the sample (i.e., at the apexes). It should be noted that no clear relationship between crack severity and crack location was identified, even though the Simufact simulations indicated there was a difference in the thermomechanical history at each location. Therefore, in order to develop a robust crack susceptibility index that accounts for both crack severity and 0204060801001201400200400600800100012001000 1500 2000Stress (MPa)Temperature (°C)Time since beginning of print (s)(a)Temperature Stress0204060801001201400200400600800100012001000 1500 2000Stress (MPa)Temperature (°C)Time since beginning of print (s)(b)Temperature Stress69 the local level of restrain on the crack, similar to the CRC test [27], modifications to the existing design are required. By example, the leg widths could be varied across the width of the sample, to create asymmetrical and varying levels of self-restraint on each of the apexes, similar to varying rod lengths in the CRC. Nevertheless, cracks were observed in two separate IN625 samples, though the quantitative measurements of crack lengths were not exceptionally repeatable between the two replicate samples, though no distinct outliers were present. However, it was found that the apex-to-tip:total crack length ratio was consistently between 0.8 to 0.9 for all apexes across both replicates. From the detailed EBSD analysis of the crack propagation, it was found that the cracks tend to follow melt pool boundaries. Therefore, this ratio may serve as a method of quality control and crack quantification for process optimization, as process parameters will affect the melt pool size and shape and, consequently, the crack length and propagation path. The design was also able to discriminate between materials, as fewer and less severe cracks were observed in IN718 compared to IN625. This result is in general agreement with prior work that has shown IN718 to be less crack susceptible than IN625 [161]. The composition effect on solidification crack susceptibility (SCS) can also be assessed for the specific powder compositions used in this study, by using the ThermoCalc data presented in Figure 33 and Kou’s SCS index [43], [169]: 𝑆𝐶𝑆 = max |𝑑𝑇𝑑√𝑓𝑠|𝑓𝑠→𝑓𝑠,𝑐𝑜 (eq. 3) in which T is temperature, fs is the solid fraction, and fs,co is the solid fraction corresponding to dendrite coalescence, which is generally assumed to occur around fs = 0.98. From the data in Figure 33, it was found that SCSIN625 = 10,089°C and SCSIN718 = 2,043°C, confirming greater susceptibility of IN625. Robust crack quantification using the current design faces a few issues, namely the presence of other defects (i.e., voids) located around the crack site, which confounds the crack morphology and size. Another issue is the varying width of the cracks, which may be due to the propagation of the crack into or out of the observed plane, suggesting that the entire crack cannot be observed in a single plane and minor differences in sample preparation can reveal or remove parts of the crack. This complicates measurement of crack lengths, as distinct locations for the end of the crack are difficult to determine with cracks that transition from wide to hairline as they grow. In summary, the existing benchmark provides a reliable method to generate cracks at predefined locations, and the crack severity can be quantified fairly easily through optical microscopy. It also seems that the dimensionless crack ratio developed here may provide a method to quantify the effect of process parameters on cracking, based on the change in melt pool geometry. The benchmark design also enables the differentiation of crack susceptibility between different alloys. However, further improvements are needed 70 for the design. In particular, the benchmark would ideally vary the extent of restraint throughout the sample to further differentiate crack susceptibility of materials and would have a robust way to deconvolute the effect of other defects that may be present in the material. However, the most important consideration is whether the results from this test are due to SC or another mechanism. 4.4.2 Crack mechanism As previously noted, several crack mechanisms can occur during solidification and cooling of Ni superalloys including SC, LC, and DDC. Each of these crack mechanisms corresponds to a particular temperature range and each typically exhibit characteristic features [44], [144]. In the case of SC, cracks occur during the terminal stages of solidification between coherency and coalescence when liquid flow becomes interdendritic and, therefore, tend to exhibit a dendritic structure on the crack surface. Moreover, it is known that SC tend to occur intergranularly, particularly at high angle grain boundaries and between coarse, columnar grains [73], [142], [170]. The current analysis of the cracks at the apexes in the benchmark sample revealed that the cracks were transgranular, resulting in a dimpled crack surface (Figure 26), and tended to have long straight sections that did not follow grain boundaries. Rather, the cracks were found to propagate through columnar grains near the melt pool boundaries (Figure 27), and a large amount of LAGB were identified along the crack surface suggesting that the regions experienced significant plastic deformation [171]. Carter et al. [163] previously noted that such straight cracks are indicative of a solid-state fracture rather than SC. The Simufact simulations also predicted that the local stress at the apexes could reach levels close to the elevated temperature tensile strength of IN625, further supporting the likelihood of a solid-state crack. The possibility of these larger transgranular cracks acting as initiation sites for SC in subsequently deposited layers was also investigated. However, the EBSD and EDS analyses in Figure 28 and Figure 30, respectively, revealed that the hairline cracks remained transgranular and did not have a strong relationship with the location of interdendritic phases (i.e., final liquid during solidification). Young et al. [172] suggested that DDC, which is a solid state crack occurring between ~0.5-0.8Tm, may occur due to the presence of carbides precipitating at the grain boundaries, and Pereira et al. [173] noted the presence of carbides rich in Nb, Mo, and Cr in Inconel 625. The EDS maps shown in Figure 30 reveal the presence of Nb, Mo, and Cr along the propagation path of these hairline fractures, and the XRD data in Figure 31 contained peaks corresponding to carbides, even though no carbon was detected in the ICP-OES analysis of the starting powder. However, DDC is also known to be an intergranular fracture mode, related to grain boundary sliding and embrittlement [174]. From the current set of results, it appears that the cracks observed within the benchmark sample are a ductile fracture that occurs when the local stress at the apexes exceeds the elevated temperature tensile strength of the material. Consequently, the existing benchmark design should be modified to reduce the localized stress at these locations. The applied stress should be 71 sufficiently high such that it could cause a shrinkage cavity of a critical size (governed by the Griffith criteria in fracture mechanics), to propagate as a crack along interdendritic films [22], but remain below the elevated temperature tensile strength of the parent material. As discussed previously, varying levels of self-restraint could be induced by varying geometric features in the benchmark, such as the width of the legs, similar to the CRC and Houldcroft tests in casting and welding. 4.5 Conclusions A lack of methods to systematically quantify solidification cracking presents a serious challenge to standardization of AM. The objective of this chapter was to develop a benchmark specimen that could induce cracks at pre-defined locations that could be easily quantified. The main conclusions of this study are as follows: 1. A SC test for AM can successfully utilise self-imposed stress (i.e., self-restraint) to promote cracking in specific regions. In the case of the current benchmark design, the apex between legs were used as stress risers or hot spots to localize stress and promote cracking at these locations. From analysis of all apexes across multiple samples, it was found that the apexes were an effective crack initiating mechanism. 2. The benchmark was able to demonstrate the effect of alloy composition on crack susceptibility, as IN718 had less cracking than IN625, which is in agreement with literature as well as solidification simulations. Although no SC was observed, the same crack morphology and propagation was seen in both IN625 and IN718, suggesting the same crack mechanism. 3. Cracks were quantified with apex-to-tip and total crack length measurements. The ratio of these two lengths showed consistency (0.8-0.9) which suggests this may be a dimensionless method of quantification. 4. No relationship between apex location and crack severity was observed, suggesting limited ability in developing a crack severity index similar to that used in CRC. 5. Cracks were found to form transgranularly and did not have a dendritic fracture surface. Moreover, found that the cracks propagated along the melt pool boundary, has a dimpled surface, and LAGB near the surface, all of which suggest that these cracks were the result of ductile fracture, not SC. Simulation results also suggest high stress at the apex locations, approaching the elevated tensile strength of IN625. 6. Simufact FE simulations produced inaccurate predictions of the temperature and stress histories, and although they confirmed that the apexes were hot spots for stress development, the absolute values have an unknown level of inaccuracy. This is mainly due to simplifications made to reduce computational demands. 72 The current specimen design presented in this chapter is an initial effort toward the development of standardized benchmarks for crack quantification. While some features of the test are promising (e.g., reliably producing cracks that can be quantified), further refinement or optimization of the design is required to induce SC and to develop more robust crack susceptibility measures. Optimisation of the design may include determination of ideal dimensions such as leg widths and heights, as well as the number of legs such that the local stress can be reliably varied throughout the sample. 73 5.0 Revised Cracking Benchmark Design 5.1 Introduction The initial benchmark design that was presented in Chapter 4, although successful at promoting cracking, exhibited several areas requiring improvement. The main issue with the benchmark design was that while cracks were observed at the apexes, they were found to not be SCs. Though this design may be useful for a general cracking benchmark, it is not suitable for a SC benchmark. Adjusting the previous design to reduce the initiation and propagation of unwanted fracture modes and instead favour SC while maintaining the core design principles that were successful (i.e., crack location, crack quantification, etc.,) may prove an efficient method of design reiteration. Furthermore, additional changes to further improve quantification of cracking can be trialed in order to achieve similar crack susceptibility indexing as seen in CRC test methods. Therefore, the objective of this chapter is to outline efforts taken to refine the crack benchmark design that was developed in Chapter 4. 5.2 Experimental methods 5.2.1 Design improvements The FE simulations in Chapter 4 show areas of high stress and strain concentration in the first layer that connects the separate legs as soon as the layer is deposited. This suggests the high stress and strain locations at the intersection of the legs serve as a good location for crack analysis. A redesign of the specimen can remove unnecessary height of the specimen to reduce material usage and process time, as the maximum length of the cracks <2mm. From the FE simulation results in chapter 4.3.5, it was also found that these leg intersections experienced very high levels of stress, which caused the formation of cracks more severe than the desired solidification cracking. Due to this, redesigning the specimen to not only reduce the level of stress, but also vary the stress at each of these intersections would prove valuable. By having varying levels of stress at each intersection, crack susceptibility could be assessed based on the number of legs experiencing cracking and the severity of cracks at different levels of stress. For example, a design that produces cracks in a low-stress intersection for a given set of parameters but produces no cracks at the same intersection under a different set of parameters can allow for clear comparison of parameter sets. Several design improvements were made to the benchmark design from Chapter 4, not only to improve the crack initiation mechanism to facilitate the initiation of SC and improve the printability of the design, but also to improve the repeatability of recorded results in order to facilitate the development of a location-based crack susceptibility index. The large tensile stress imposed at the apex of the leg intersection was great enough to cause transgranular fractures and, as such, should be reduced to avoid these fractures, but large enough to promote SC. To achieve this, small radii were added to the apexes between the legs in order to reduce the stress concentration [175]. Additionally, the width of each leg was varied linearly from one 74 end of the benchmark to the other, so that different levels of restraint could be imposed on each apex, similar to the concept of the vary lengths of the rods in CRC and the varying length of slits in the Houldcroft test [25], [27]. This change is intended to impose different levels of stress at each apex so that a similar quantification method to that of CRC can be utilized and a normalized index of crack susceptibility can be developed. The angle between each leg was also reduced in order to minimize the overhang as the sample is being printed. This was intended to improve printability for lower resolution processes such as DED (compared to processes such as PBF). This change ultimately leads to a greater height of each leg, and thus the entire sample, and due to this, the initial vertical sections of the legs have been removed so that the inclination begins immediately as deposition of the second layer begins. A schematic diagram of the revised benchmark design can be seen in Figure 36. Figure 36. Design schematic of a potential redesign with a thickness of 10mm. All units are in mm. 5.2.2 Fabrication Fabrication of the new design was performed on a Liburdi Automated Welding System (LAWS) 250 with a spot size of 0.3 mm. The same scan strategy used for the original design was implemented for the redesign (Figure 22), and samples were printed using gas-atomised Inconel 625 powder supplied by Carpenter Additive. Two replicate samples were printed for repeatability analysis. The laser power was constant throughout all layers of this benchmark design, unlike in the prior benchmark design, and all process parameters can be seen in Table 10. 75 Table 10. L-DED process parameters Process parameter (units) Value Powder feed rate (g/min) 4.0 Travel speed (mm/min) 1000 Layer height (mm) 0.3 Ar shielding gas flow rate (L/min) 11.8 Ar nozzle gas flow rate (L/min) 2.8 Ar powder carrier gas flow rate (L/min) 3.5 5.3 Results and discussion The refined benchmark samples were printed and can be seen in Figure 37. However, it can be seen from Figure 37 that, while the samples printed, there were severe geometric flaws during printing. The varying widths of the legs produced errors in slicing the benchmark CAD model, as the slicer software was unable to produce a toolpath that consistently increased the width of each leg with a finite weld bead width, and as a result, some layers had toolpaths that were not optimal for the desired width (Figure 38). The effect of this issue can be seen in Figure 37 where the largest leg has a clear slope on its outermost edge that was repeated for both samples. Moreover, L-DED parameters that were not fully optimized for this geometry led to variations in material deposition during printing, which can be seen in the ‘waviness’ of the deposited layers that are particularly evident on replicate 1. These issues further highlight the importance of designing for AM in developing a benchmark design, as less geometrically accurate processes (i.e., DED) may struggle with producing consistent samples. For example, the values of the print parameters greatly affects the geometry of sliced layers, as a variation in bead width and layer height will affect the detail of fine geometric features such as the small radii located at the intersection of the legs; whether these features are fully defined by the GCode varies greatly based on the defined resolution, and lower resolution processes may produce a bead that is greater than the radius of the fillet, effectively removing it entirely. It is also difficult to determine if the radii successfully formed at each apex in the printed samples, and the minimum gap spacing between two geometric features is an important consideration for AM processes. Despite printability issues, crack assessment in these samples was possible. 76 Figure 37. Initial prints of the redesigned benchmark. 77 Figure 38. (a) Demonstration of slicing issues along the height of the sample due to increasing leg thicknessesalong the length of the benchmark design. (b) Although a previous layer has successful overlap between perimeter tracks, (c) the subsequent layer has insufficient overlap. This variation between two layers of the same scan orientation is due to the direction of scanning along the axis (negative to positive directionality vs. positive to negative directionality). Initial optical microscopy of the apexes in the revised benchmark (Figure 39) shows cracks are once again present at several apexes. However, the cracks in the revised benchmark appear to be considerably less severe than those observed in the original design. There is no clear indication of the large, transgranular fractures with the zig-zag propagation path, however lack of fusion is once again present above the majority of the apexes. The lack of fusion is attributed to sub-optimal process parameters, irregularities caused by toolpath and slicing issues (Figure 38), and the need to further improve the benchmark design for printability (i.e., study effect of apex overhang angle on defect formation). Unfortunately, despite varying the width of legs throughout the design, the severity of cracking based on the apex location is not clear from these initial images. However, apex 5 for both replicate 1 and replicate 2 appears to only experience lack of fusion without any cracking. Comparatively, apexes 1-3 experience multiple defect modes, as seen in Figure 40. Currently, only optical microscopy has been performed on these samples and, as such, the detailed mechanisms of these defects have not yet been fully explored. Successful formation of the radii at 78 each apex also varies across the samples, as apexes 1 and 2 in replicate 2 appear to form the desired radius, whereas the same apexes in replicate 1 do not appear to have successfully produced the desired radii. There does not appear to be consistent geometric accuracy across apexes for both samples, and factors influencing successful radii formation cannot be determined from the initial images. These findings further emphasize the need for strong understanding of the geometric limitations of the particular printer in the development of a benchmark specimen. Figure 39. Optical images of each apex in replicates 1 and 2 of the redesigned benchmark. 79 Figure 40. High magnification optical images of cracks present in (a) replicate 1 apex 2, (b) replicate 1 apex 3, (c) replicate 2 apex 1, and (d) replicate 2 apex 3. 5.4 Conclusion Initial results indicate that the redesigned benchmark specimen can successfully eliminate the large transgranular fractures that were common in the original design. Lack of fusion is still present above the majority of apexes in the redesigned samples, however the apex with the thinnest legs (apex 5) showed no clear signs of cracking in either replicate sample. Conversely, the presence of various crack morphologies across the other apexes suggests that crack susceptibility at the apexes may be impacted by the width of the joining legs. If this is indeed accurate, a location-based susceptibility index could be developed to take this into consideration. Further numerical simulations and microstructure analysis is required to determine the severity of cracking at each apex, as well as the fracture mode of each crack, as whether or not these are solidification cracks cannot be determined with optical imaging. 80 6.0 Discussion The research presented in this thesis outlines the need for standardised testing methods for quantifying SC in AM processes and presents a first step towards this goal. The clear need for repeatable methods of testing and result quantification has been addressed as a key knowledge gap in the field and existing crack tests and AM geometric benchmarks serve as a guide on adapting existing ideas to AM-specific requirements. The presented benchmark specimens were developed with multiple AM processes and materials in mind in order to provide a robust framework for future development of SC tests for AM. However, the need for future improvements in the design were evident, as the presented benchmark was unable to produce SC that could be clearly identified and quantified. Nevertheless, the underlying design principles and benchmark designs provide valuable insight into potential methods of crack testing for AM, whether it is for SC or other crack initiation modes. 6.1 Crack formation and mechanism The initial benchmark presented in Chapter 4 was successful in its aim to reliably create cracks at specific pre-defined locations. These cracks, however, were determined to be solid-state fractures and not SCs as intended. The formation of cracks at desired locations itself serves as a valuable guide for future design iterations and this concept can be adapted for future SC tests for AM. The simplicity of the design and the self-restraint induced stresses serve as a foundation for crack-specific test development. In order to be effective as a SC initiation site, the local stress must be less than the elevated temperature strength of the material to avoid fracture, but high enough such that the stress could cause a shrinkage cavity in the mushy zone to propagate as a crack. The critical cavity size and corresponding level of stress necessary to initiate a crack are governed by the Griffith criteria [22]. The design revision that was investigated in Chapter 5, in which the apexes of the legs were rounded, was intended to reduce the local stress at the apex compared to the initial design [175]. This design revision appears to have successfully reduced stress concentration enough to eliminate transgranular fracture modes, however as the benchmark uses geometry alone to create stresses, equivalent levels of stress across different material systems may produce extensive variation in results. By example, if the benchmark geometry successfully induces a level of stress sufficient to induce SC but not ductile fracture in nickel alloys, that same level of stress in an aluminum alloy may be well beyond the range of desired stress to induce SC. To this end, test development requires an understanding of the thermomechanical properties of the alloys under consideration, as well as the expected thermomechanical history during printing. This prediction of the thermomechanical history for different materials can be improved using FE simulations, however, as seen in Chapter 4, these complex simulations require detailed fine-tuning and computationally intensive processing in order to produce accurate results. 81 6.2 Quantification and standardisation The formation of cracks in specific pre-defined locations with repeatable results across multiple samples and samples printed with different alloys, as presented in Chapter 4, demonstrates the effectiveness of this design in its aim to develop a standardised method of presenting crack susceptibility. The ability to consistently form observable cracks at pre-defined locations is an important step to quantifying crack severity and generating statistically significant results. As the design was unable to produce SC, no quantification for this specific crack initiation mode could be developed, however initial attempts were made to quantify the cracking that did occur. Based on methods presented in Chapter 2, simple crack quantification methods can be adapted for AM-specific tests once the nature of SC in such a benchmark is better understood. Chapter 2.2.4 details the different methods of SC quantification in AM, and these methods can guide future quantification methods once a benchmark design can accurately develop SC in a repeatable manner. Whether the method of quantification is based on crack length, surface area, or a SC susceptibility index remains unknown as the nature of repeatable SC in AM was not successfully demonstrated by this research, however a significant knowledge base exists for both SC testing and AM benchmarking, as demonstrated in Chapter 2. A list of criteria for an AM-specific SC test or benchmark specimen were outlined in Chapter 2.3.1. It was found that no existing cracking or welding tests were easily translated to AM, although some of the underlying design philosophies from existing tests could be adopted, and that a novel approach would need to be taken for AM. Chapters 4 and 5 presented the current efforts to develop a first of a kind SC benchmark design for AM. In this chapter, the outcomes of these benchmark designs are evaluated against the list of criteria from chapter 2.3.1. Standardisation: The results obtained in Chapter 4 provide a foundation for standardised quantification of results. Crack formation at pre-defined locations with easily observable crack propagation allowed for ease of quantification using equipment that is readily available (e.g., optical imaging as opposed to SEM imaging). This allows for greater access to a testing method that doesn’t require investment in expensive analytical equipment, aiding standardisation and industry adoption. While a crack susceptibility index similar to that used for the CRC test was not able to be developed, analysis of the cracks in the AM benchmark revealed a constant ratio between the apex-to-tip length (linear) and total crack length (zig zag) of 0.8-0.9. Analysis of the crack propagation path revealed that the crack tended to follow melt pool boundaries and, therefore, this ratio could potentially be used to optimize process parameters, since the selection of process parameters affects the melt pool size [176]. Simple yet robust design: The designs used relatively simple geometries to induce self-restraint on the solidifying metal. This simplicity improves the applicability of the design to multiple AM processes and 82 reduces material cost due to the small size. Despite the relatively simple geometric design, the benchmark was effective at producing cracks at desired locations and varying the level of cracking in different materials (IN625 and IN718), making it fairly robust. However, while the designs were relatively simple geometrically, these were not necessarily simple designs for printing, particularly for L-DED. Lower resolution processes (such as DED) may struggle to generate certain design features (e.g., the radii at the top of each apex in Chapter 5). Moreover, despite the relatively simple geometry of the design, numerical simulations with commercially available FE software were complex and computationally expensive. Thus, the design and printability of the benchmark design relied heavily on trial and error which proves to be very costly for AM processes. Attempts to reduce the complexity of the simulations severely limited the accuracy of results. Assessing multiple parameters: Results from Chapter 4 were gathered from a single parameter set with little to no variation and as such an analysis of the effect of different parameters was not performed. With that in mind, simulation results show that thermal history plays an importantrole on the development of stress at each apex. As such, it could be expected that variation in parameters, and thus thermal history, would impact the stress development and thus crack formation at each apex. Parameter selection also affects the resolution of the print, and despite process parameters being optimised for simple geometries such as cubes, simple variations in geometry (e.g., varying leg thicknesses) can cause issues with tool path generation which in turn results in poor geometric fidelity as described in Chapter 5. This poor geometric accuracy can then result in geometric and metallurgical defects that confound crack analysis. To this end, the development of a crack benchmark requires a substantial investment in design for AM to account for these issues, and this varies greatly between processes such as DED and PBF. The ability of the test to evaluate a wide range of parameters is thus limited by the effect of parameters on geometric accuracy. Assessing alloy composition: The effect of alloy composition was investigated with the comparison of Inconel 625 and Inconel 718 in Chapter 4. The results from this investigation showed more severe cracking in Inconel 625 compared to Inconel 718, and this is consistent with previous studies of Ni-based superalloys and thermodynamic simulations of solidification (Chapter 4.3.4). This indicates that the design can distinguish between alloys of the same family. However, the ability of the design to evaluate differences in crack susceptibility for alloys in different classes (e.g., Al and Ni) cannot be determined from the current results. Applicability for multiple vendors: Again, the simple design of the benchmark was intended to allow for printers of different capabilities to produce a sample. This has not been confirmed in this thesis, due to time and equipment availability constraints, so the exact range of machines that are able to reproduce this design is unknown. However, it should be noted that the original benchmark was printed on a LAWS 1000 printer, 83 while the redesigned benchmark was printed on a LAWS 250 printer. While there were some challenges with the printability of the redesigned specimen, it is reasonable to assume that the benchmark design could be printed on other L-DED systems. Applicability for multiple AM processes: Applicability to other processes is once again impacted by the printability of the design. As the benchmark is designed to be printed without supports, this allows for application with DED processes as well as other processes such as PBF or SLS. The effectiveness of this design with other processes has not been determined in this thesis, due to time and equipment availability constraints, and is an area for future study. The evaluation of transferability across AM processes is outside the scope of the current research, however it should be noted that processes such as L-PBF are expected to be able to produce this design more effectively, given the better geometric resolution of the process and the fact that unfused powder in prior layers provides some support to features such as overhangs. Therefore, the effectiveness of the designs presented in the current research, although promising, have not been fully realised in this study and several major issues have been observed. Other than the issues such as inability to produce isolated SC and the impact of scan strategy that have already been discussed, the design itself has not been optimised for material consumption. It has not been determined how many leg intersections are necessary, the ideal width of the legs, or the ideal height of the sample. These are all aspects of the design that may require improvement but have not been investigated in this thesis. Ideally, future designs would be guided by numerical approaches to reduce trial and error, however the current state of FE simulations reduce the effectiveness of this approach due to the high computational demands. Further improvements in computational software will aid in design and optimisation of benchmarks. Nevertheless, the results from the initial benchmark designs are promising, as the reliable and repeatable formation of cracks in specific locations has not been previously demonstrated. Result quantification and effects of alloy composition were also demonstrated by this design, providing a good starting point for future work in the standardisation of AM testing. 84 7.0 Conclusions and Future Research 7.1 Summary and conclusions The research presented in this thesis serves as a novel investigation into the development of a SC test for AM. Using the design philosophies of existing tests in casting (CRC) and welding (Houldcroft) that are well established provides an important first step in designing for AM, and existing geometric and mechanical AM benchmarks help guide AM-specific design. The utilisation of self-induced restraint to generate stress localization and promote cracking has been seen to effectively produce repeatable results in casting and welding, and this concept was the foundation for the development of an AM-specific test which can take advantage of more complex geometries to impose this self-induced stress. Although a foundation can be provided by existing tests, AM-specific challenges require a unique approach to fully encapsulate the complexities of these processes while not compromising ease of fabrication and quantification of results. As no existing tests for SC exist for AM, initial designs are unlikely to fully capture these complexities, but they can provide important results to base future tests off, such as tests specifically focussing on a single AM process or material class. The designs presented in this thesis are a promising first step in the development of future SC tests for AM, but further work is required to refine the designs to a point that they could be widely used by researchers and practitioners. The main conclusions of this thesis are as follows: 1. No tests exist to systematically evaluate SC in AM. Existing tests for casting and welding highlight the usefulness of self-restraint tests, and these types of tests are well-suited for AM processes. However, there is no existing SC susceptibility test from casting or welding that is well suited to translate to AM. The track-by-track and layer-by-layer nature of AM presents a unique thermomechanical history and possibility for highly complex geometries, and these effects could not be captured by existing tests. However, core design philosophies, such as the use of self-restraint to impose localised stresses, proved effective in generating crack in the AM benchmark. 2. A novel cracking benchmark design was developed specifically for AM. The proposed benchmark design was able to produce cracks successfully and repeatedly at pre-defined locations. This allowed for cracks to readily be identified and measured across multiple samples. This is an essential feature of a reliable standard. Moreover, measurement and analysis of cracks revealed that there was a fairly constant value of the ratio of apex to tip: total crack length, and microstructure analysis suggested that this ratio could be used to assess the effect of process parameters on cracking due to the relationship between parameters and melt pool size. 3. No relationship was found between crack severity and crack location (i.e., apex number) within the benchmark samples. Moreover, due to computational limitations, the thermomechanical history and 85 local stress at each apex has an unknown level of prediction error. As a result, it was not possible to develop a normalized crack index as has been developed in the past for other crack tests such as the CRC. 4. While cracks were observed in the AM benchmark, they were found to not be SC. Microstructure analysis revealed that they were transgranular cracks and likely ductile fracture of the material at elevated temperatures. Future benchmark designs would require the ability to fine tune the local stress to be less than the elevated temperature tensile stress of the material, but high enough that cavities in the mushy zone could propagate as a SC. 5. Numerical FE simulations can provide valuable insight into the effects of scan strategy, bead width, and layer height, as well accurately predict regions of localised stress. However, due to the computational intensity of simulating printing of complex geometries, the predicted thermomechanical values (e.g., temperature, stress, and strain) have an unknown, but significant, margin of error. 6. Design revisions demonstrate the impact of geometry in stress development, as well as the influence of process parameters tool path generation on geometrical accuracy. A substantial effort towards design for AM is required to understand the optimal features and dimensions for a given process, as lower resolution processes, such as DED, may struggle to generate small features such as the radii at each apex in the redesigned benchmark. Nevertheless, the revised benchmark design was able to produce cracks at pre-defined locations, and preliminary analysis suggests that varying leg widths may be an effective method to vary the stress and crack severity at different apex locations in the sample. The ultimate goal of this research was not to develop the final SC test method for AM, but rather to provide an initial benchmark that can serve as useful guides for further development. To this end, the research was successful in that it provides a novel assessment of the current state of SC research in AM and presents a path forward for an important field of study of one of the most significant defects in AM. 7.2 Limitations of the current study The research presented in this thesis presents a number of interesting findings, however certain limitations can affect the quality and reliability of results. Being aware of these limitations is important for future research to ensure high quality results. One limitation is the geometrical accuracy of the printer or process. Accurately capturing the desired geometries, especially those with dimensions of only a few millimetres, is very important for a test which relies on the benchmark geometry to induce the desired effects. L-DED suffers from poor surface finish and geometric accuracy compared to processes such as PBF or sintering. Being aware of this limitation can help guide development for a test that is designed for a specific process. 86 Another limitation is the reliance on one observation plane to record measurements. It was hypothesized that cracks may propagate into, or out of, the observed plane, which obfuscates the true length of the fracture. The observed plane is highly dependent on sample preparation techniques, as even a deviation of less than 1mm may result in a crack being hidden from view. The magnitude of this limitation is difficult to quantify with a small sample size, however evidence is present as seen in the hairline fractures depicted in Figure 30 that may be a transgranular fracture that is propagating out of the observed plane. The need for optimised parameters to reduce porosity may also present complications if the parameters to be tested cover a wide range of values. That is, if test parameters induce a greater level of porosity, these voids may obfuscate the true crack dimensions and hinder result quantification. The requirement for relatively high-density manufacturing may impact some desirable research. 7.3 Future research There remains a significant amount of research necessary for a fully realized AM-specific SC susceptibility test and many potential directions that this research can take. Further improvement of the presented benchmark may yield reliable results, or a complete redesign using only the core concepts of the benchmark may be more beneficial. Comparing results across different AM processes, material families, and parameter sets is needed to confirm the consistency of results and to evaluate the range of applicability of the presented design. The following are recommended as next steps in the development of a robust SC design for AM: 1. Evaluate the geometric limitations of a given AM process and even the specific printer model that is used to fabricate the benchmark design. Effective analysis of SC susceptibility is best focussed on a singular process for initial investigation, as the wide range of limitations across different processes make it difficult to account for each process effectively. By focussing effort into a singular process, a design that can accurately test SC susceptibility can be developed, and subsequent efforts to adapt this design for other processes can follow once consistent results are obtained. 2. Reduce the effect of variation in scan strategy and layer height. This can be achieved either through design alterations or defined process parameters. As varying process parameters is a desired area of study, significant constraints placed on the range of parameters would be counter-productive, and thus the optimal approach would be design-based. However, this is more difficult to implement, as accounting for all possibilities is beyond the scope of any one test, and thus an optimal balance between the two approaches is necessary. 3. A design which can implement a SC susceptibility index into the analysis of results is ideal, however this index does not necessarily require the location-based method discussed in this thesis. 87 Although this approach has been proven successful in past tests (namely, the CRC test), alternative approaches might be better suited for AM processes. 4. Once repeatable SC initiation has been achieved, ideal result quantification would not rely on consistent analysis of crack morphology in order to confirm the nature of crack initiation. A test that can produce SC and can be assessed with minimal investment in analytical equipment is desirable for widespread adoption. 5. The combination of a SC susceptibility test with a related mechanical test that can evaluate the effect of SC on mechanical performance may serve as the ideal test procedure, as although SC susceptibility may be determined from one test, the effect of SC is not demonstrated. A mechanical performance test designed specifically to compliment a SC susceptibility can provide a well-rounded evaluation. Once a design successfully produces repeatable SC at a desired level, the next key step is the standardization of result quantification. Assessing which factors of the crack (location, size, shape, etc.,) are the most significant identifiers of crack susceptibility is needed to develop a consistent method of analysis that can be used as a standard method of SC. This method of quantification will vary based on the nature of the test, and a universal SC susceptibility quantification method for AM is unlikely, especially given the large variety of AM-based processes. 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Thesis/Dissertation
eng
Mechanical Engineering
Vancouver : University of British Columbia Library
Attribution-NonCommercial-NoDerivatives 4.0 International
Graduate
Benoit, Michael J.
On the development of a novel solidification crack test for additive manufacturing
Text