Analysis of single stage SPIF process applied to the hole flanging operation.

AnalysisofsinglestageSPIFprocess appliedtotheholeflangingoperation DPI-2012-32913 Doctoral Thesis MarcosBorregoPuche Analysisofsinglestage SPIFprocessappliedto theholeflanging operation Marcos Borrego Puche 2018 Doctoral Thesis Analysis of single stage SPIF process applied to the hole flanging operation Author: Marcos Borrego Puche Mechanical Engineer Directors: Dr. Carpóforo Vallellano Martín Dr. Domingo Morales Palma Dpto. de Ingenieria Mecánica y de Fabricación Escuela Técnica Superior de Ingeniería Avda. De Los Descubrimientos s/n. 41092 Universidad de Sevilla Sevilla, 2018 ii iii A mi familia iv v A CKNOWLEDGEMENT Parecia que nunca iba a llegar el momento de los agradecimientos y sin embargo el momento ha llegado. Quiero agradecer en primer lugar a los doctores Prof. Carpóforo Vallellano y Domingo Morales por su implicación en este proyecto, tanto en la dirección como en las correcciones. Al Dr. Andrés Martínez-Donaire, disponible 24/7 para cualquier cuestión. Al Dr. Gabriel Centeno, guerrillero urbano que se las sabe todas. A Antonio Guillén, buena persona y mejor amigo, con el que disfruto enormemente en el taller. A Manuel Sánchez, excelente ganadero ecológico y miembro secreto del equipo “ARUS”. Al Dr. Alfredo Navarro, que poco a poco vamos a cambiar la ingeniería por la botánica. A José el carpintero, excelente profesional y compañero. Me encanta charlar con él. A Juan Martínez, servicial y gran alumno de Solidworks. A Manuel de fluidos, que tiene un coche con un catalizador que no se lo merece…. Al futuro Dr.Andrés Lopez, que como a mi, le sube la adrenalina con lo que otros tiran. Al grupo del Técnico de Lisboa, por su trato acogedor y familiar: Prof. Dr. Paulo Martins, visto sus “ferramentas” en otra vida seguro que fue discípulo de Leonardo Da Vinci. Dra. Beatriz Silva, diligente y dispuesta a reir en todo momento, amante del queso y el vino. Dr. Carlos Silva, no me lo ha confirmado pero creo que es familia del Sr. Gomes… Dr. Luis Alves, excelente persona y potencial artesano de cascavéis de falcoaria. Dr. Valentino Cristino, servicial y gran conversador portugués de ultra mar. A las secretarias Paula y Fatima, cordiales, serviciales y familiares. A mis padres Antonio y Ana, los impulsores sin ninguna duda de esta aventura, lo vieron claro desde el principio y animaron sin fisuras. Sólo ellos son conscientes de lo que disfruto con ellos en su casa, también la mia. A mis hermanos Pablo y Berta, iniciadaos antes que yo y precursores de mi carrera docente. Les deseo lo mejor. No puedo olvidar a mi tia Lola, doctorada de la vida con casi 80 años de experiencia recientemente fallecida. Cita ineludible de cada miércoles. Y por supuesto le doy las gracias a Patricia por su “pacencia”” infinita”. Alguno se me olvida, que no me lo tenga en cuenta,¡gracias! vi vii R ESUMEN Tradicionalmente, las operaciones de rebordeado se llevan a cabo en la industria mediante técnicas de prensado convencionales utilizando punzón y troquel. Sin embargo, investigaciones recientes en conformado incremental (ISF) han señalado ventajas significativas de esta operación con respecto a los procesos convencionales. Entre otras son, la flexibilidad, la relación coste-efectividad y una importante mejora aparente en la confomabilidad del material. En este sentido, la presente tesis doctoral ofrece un análisis crítico de la operación de rebordeado de agujero cilíndrico mediante el conformado incremental mono punto (SPIF), realizada en una etapa, desde el punto de vista de la conformabilidad del material y las capacidades geométricas del proceso. Para comprender y cuantificar las ventajas de las operaciones de SPIF de una sola etapa, se llevó a cabo una campaña experimental de rebordeado de orificios cilíndricos mediante conformado incremental y convencional. Esto se realizó sobre una lámina de aleación de aluminio 7075-O. Se proporciona un estudio detallado de la evolución de las deformaciones en el flanco. Éstos se midieron utilizando métodos avanzados de videogrametría y fotogrametría. Los mecanismos y modos de fallo de la lámina y la caracterización de la capacidad de rebordeado en ambas operaciones se analizan usando el diagrama límite de conformado (FLD). El efecto de flexión inducido por la herramienta en SPIF y el radio de esquina en el caso convencional ha sido claramente dilucidado, junto con la geometría del flanco y las fuerzas de conformado involucradas. Finalmente, la comparación llevada a cabo permite extraer las ventajas y limitaciones tecnológicas en el proceso de rebordeado de orificios mediante SPIF de una sola etapa, y descubrir futuras líneas de investigación para mejorar su aplicación práctica. viii xv   Equivalent stress t Thickness t o Initial thickness   Average thickness t o /R Bending ratio V Volts Y Yield stress Δz Step down xvi xvii L IST OF ACRONYMS AA1XXX-YZ Aluminium alloy, series 1(e.g. AA1050-H14) AA2XXX-YZ Aluminium alloy, series 2 AA3XXX-YZ Aluminium alloy, series 3 AA5XXX-YZ Aluminium alloy, series 5 AA6XXX-YZ Aluminium alloy, series 6 AA7XXX-YZ Aluminium alloy, series 7 ACW Anti-Clockwise ASTM American Society for Testing and Materials CAD-CAM Computer-Aided Design/ Computer-Aided Manufacturing CCD Charge-Coupled Device CIRP The International Institution for Production Engineering esearch CNC Computer numerical control CW Clockwise DIC Digital Image Correlation EP European Patent FEA Finit Element Analysis FFL Fracture Forming Limit FLD Forming Limit Diagram FLD0 FLC under plane strain conditions FLD0incr FLC under plane strain incremental conditions H2O Water xviii HCl Hydrochloric acid HER Hole Expansion Ratio HFl Hydrofluoric acid HNO3 Nitric acid HRC Rockwell C Hardness ISF Incremental Sheet Forming LFR Limiting Forming Ratio ORB Orbital Roller Ball PTFE Polytetrafluoroethylene SFFL Shear Fracture Forming Limit SPIF Single Point Incremental Forming SPIF-A Single Point Incremental Forming-Aveiro TPIF Two Point Incremental Forming TR Average thickness ratio UNE Una Norma Española US United States UTS Ultimate Tensile Streght VRB Vertical Roller Ball xix I NDEX OF F IGURES Figure 0.1 Thesis structure .......................................................................................... xxxi Figure 1.1 Example of flanging by conventional press working (Custompart.net) ........ 1 Figure 1.2 Types of conventional flanges: A) straight, B) stretch, C) shrink, D) reverse, E) Joggled and F) hole-flanging ........................................................................... 2 Figure 1.3 Schematic of conventional hole flanging, (left) specimen with an initial precut hole and finished hole flanging, and (right) the conventional press working setup. Montanari et al.,2013 ........................................................................................................ 3 Figure 1.4 Model (left), punch load vs. displacement (centre) and strain paths at the flange tip for different punch shapes (right). Tang ,1981. ................................................ 4 Figure 1.5 Schema of the testing setup (left) and detail of specimens with a blank holding effect (right). Krichen et al.,2011 ......................................................................... 5 Figure 1.6 Schema of the conventional hole flanging process with clearance indication. Kacem et al.,2011 ............................................................................................................... 6 Figure 1.7 FE and experimental shape of flanges for different R c. Kacem et al.,2011 ...... 6 Figure 1.8 Flange length and thinning evolution for different R c. Kacem et al.,2011 ..... 7 Figure 1.9 Distribution of the stress triaxiality in the conventional hole flanging process on aluminium A1050-H14 with and without ironing. Kacem et al.,2013 ......... 7 Figure 1.10 Failure map for the conventional hole flanging process at a fixed punch diameter. Huang et al.,2001 .............................................................................................. 8 Figure 1.11 Failure map for the conventional hole flanging process based on process and geometrical parameters. Soussi et al.,2016 ................................................................ 9 Figure 1.12 Relation between the diameter of the expanded hole and diameter of the initial hole for various punch profile radii. Huang et al., 2001 ...................................... 10 Figure 1.13 Simulated relations between the diameter of the expanded hole and xx diameter of the initial hole for various truncated conical punches. Huang et al.,2002 . 10 Figure 1.14 Image of apparatus and process for incremental dieless forming. Patent US3342051A ................................................................................................................... 11 Figure 1.15 Method of dieless forming surfaces of revolution. Patent US 3316745 ..... 12 Figure 1.16 Illustration of SPIF a) and TPIF b) and c) with partial and full die. Jeswiet et al.,2005 ............................................................................................................ 13 Figure 1.17 Patent by Tuominen of Twincam OY, Finland EP 1560668 .................... 14 Figure 1.18 Concept of Zhang with three robots. Zhang,2007 ...................................... 14 Figure 1.19 Patent number [EP 1899089] from Univ. Leuven 2005 ........................... 15 Figure 1.20 Formed by SPIF headlights and manifolds by Jeswiet et al.,2001 .............. 16 Figure 1.21 Service panels before and after SPIF forming by Amino et al.,2002, 2014 16 Figure 1.22 Cranial plate capture (left) and SPIF cranial plate (right) by Bagudanch et al.,2015 ......................................................................................................................... 16 Figure 1.23 CAD model (left) and built machine (right) of SPIF-A. Alves de Sousa et al.,2014 ............................................................................................................................. 17 Figure 1.24 Geometries (left) of tests used to calculate (right) SFFL and FFL in AA1051-H111, isik et al.,2014 ....................................................................................... 18 Figure 1.25 Fracture localization due to the bending effect in AA1100. Fang et al.,2014 ............................................................................................................................. 19 Figure 1.26 Percentage contribution of each deformation mode for four elements under varying wall angles. Maqbool et al.,2017 ............................................................. 19 Figure 1.27 Percentage contribution of each deformation mode for two elements under the influence of the a) tool diameter, b) thickness, c) friction and d) step down. Maqbool et al.,2017 .......................................................................................................... 20 Figure 1.28 Schematic of the Law of Cosines. Hussain et al.,2007 ................................ 22 Figure 1.29 Angle, flat, parabolic and hemispherical tool shapes tested by Cawley et al.,2013 ............................................................................................................................. 23 Figure 1.30 a) Rigid hemispherical tool, b) vertical roller ball and c) oblique roller xxi ball. Lu et al.,2014 ............................................................................................................ 23 Figure 1.31 a) Truncate cone geometry of surface roughness tests, b) part processed with a rigid tool, c) part processed with VRB and d) with ORB. Lu et al.,2014 ........... 24 Figure 1.32 Result of surface roughness tests in terms of Rz and Ra. Lu et al.,2014 ... 24 Figure 1.33 Roughness (left) and temperature increment (right) for different levels of tool rotation (rpm). Durante et al.,2009 ......................................................................... 25 Figure 1.34 Laser texturized surface tool head. Xu et al.,2013 ...................................... 26 Figure 1.35 SPIF machine with dynamic laser supported heating and formability tests produces 0.6 mm TiAl6V4 sheet with and without heating. Duflou et al.,2008 .......... 27 Figure 1.36 Electric heating system proposed by Fan et al.,2008 .................................. 27 Figure 1.37 Max wall angle for different feed rates and materials. Bastos et al.,2016 .. 28 Figure 1.38 Values of Ra for four materials and feed rates. Bastos et al.,2016 .............. 29 Figure 1.39 Schema of dieless incremental forming. Jeswiet et al.,2005 ........................ 29 Figure 1.40 Basic tool path strategies: left) z-level and right) helical path for pyramid. Skjoedt et al.,2007 ............................................................................................................ 30 Figure 1.41 70 O truncated cone in AA1050 for single-stage strategy, a) conventional three-stage strategy b) and improved three-stage strategy by sequential limit analysis c). Mirnia et al., 2014 ...................................................................................................... 30 Figure 1.42 Tool path strategies and corresponding thickness distributions for an initial pre cut hole with d o =35.5 mm and final hole diameter of d f =65 mm in AA1060. Cui et al.,2010 .................................................................................................................. 32 Figure 1.43 a)Experimental strains in conical hole flanging with multi-stage strategies in AA1050 b) and selected steps in the successful production of cylindrical hole flanging. Centeno et al., 2012 .................................................................................. 32 Figure 1.44 Strain distributions along meridional directions of hole-flanged parts in a) AA1050-H111 and b)titanium (grade 2) under press working conditions and multistage SPIF. Silva et al.,2013 ............................................................................................ 33 Figure 1.45 High-speed incremental hole flanging: tool designs a), setup b), parameters c) and specimens before and after performing. Bambach et al.2014............ 34 xxii Figure 1.46 Different tool path strategies, forming forces and timing for both tools. Cao et al., 2016 ................................................................................................................. 34 Figure 2.1 Geometry of the tensile test specimen and strain-stress curves in three different directions ........................................................................................................... 36 Figure 2.2 Section of the bulge test setup (recreation) ................................................... 38 Figure 2.3 Configuration of cameras for the bulge test under the video correlation system ............................................................................................................................... 39 Figure 2.4 Equivalent stress vs equivalent strain for tensile and bulge tests ................ 39 Figure 2.5 Stack compression test setup ......................................................................... 40 Figure 2.6 Equivalent stress vs. equivalent strain for tensile, bulge and stack compression tests ............................................................................................................. 41 Figure 2.7 Forming Limit Diagram for a ductile material AISI 304-H111. Centeno et al., 2017 ............................................................................................................................ 42 Figure 2.8 Recreation of Nakazima test (section of setup) ............................................. 42 Figure 2.9 Geometry of uniaxial, equibiaxial, near plane strain and biaxial strain Nakazima tests ................................................................................................................. 43 Figure 2.10 Setup of Nakazina tests ............................................................................... 44 Figure 2.11 Conventional Formability limits at necking (FLC - Forming Limit Curve) and fracture (FFL - Fracture Forming Line) for AA7075-O metal sheets with a thickness of 1.6 mm. Forming limit strains in stretch-bending tests near plane strain for cylindrical punches with radii of 10 and 6 mm. Borrego et al., 2015 ....................... 45 Figure 2.12 Schema of the stretch-bending test, R6 and R10 punches .......................... 46 Figure 3.1 Conventional hole flanging process scheme .................................................. 47 Figure 3.2 Set of punches with different edge radii for hole flanging tests by conventional press working (6 mm, 8 mm, 10 mm and 47.9 mm) ................................ 48 Figure 3.3 Electro etching setup ..................................................................................... 50 Figure 3.4 Examples of black on white and white on black markings ........................... 50 Figure 3.5 Specimen before and after ARGUS computation; an example in xxiii conventional and SPIF processes .................................................................................... 51 Figure 3.6 Flanges from pre cut hole diameters of 59 (fractured), 60 (necked) and 61 mm (successful) with a 6 mm edge radius ...................................................................... 52 Figure 3.7 Non-dimensional flange height (h/d f ) of successful conventional holeflanged sheets for different pre cut diameters .................................................................. 55 Figure 3.8 In-plane strains along the meridional direction for all successful hole flanged parts fabricated using different punches ............................................................ 56 Figure 3.9 In-plane strains along the meridional direction for hole-flanged parts using pre cut holes with a 61mm (successful) and 60mm (necked) diameter and different punch edge radii. Comparison with the results for the stretch-bending tests using cylindrical punches with a 6 and 10mm radius and the extrapolated FLC with the bending contribution ....................................................................................................... 58 Figure3.10 In-plane strain paths during the hole flanging process obtained from the interrupted tests using the pre cut 61 mm hole and a 6 mm punch edge radius .......... 59 Figure 3.11 Vertical forming forces for a 65 mm diameter pre cut hole and punch corner radii of 8 mm and 6 mm ....................................................................................... 61 Figure 3.12 Section of specimen successfully conformed with a R=8 mm corner radius punch and 61 mm pre cut hole, which was scanned and compared with the theoretical profile ................................................................................................................................ 62 Figure 4.1 (a) Experimental setup for SPIF. (b) Scheme of the single-stage hole flanging process by SPIF. (c) Scheme of the z-level trajectory for the conical frustum and cylindrical helical path .............................................................................................. 66 Figure 4.2 Forming tools used for SPIF ......................................................................... 66 Figure 4.3 Pictures of the closest success-failed test pair with a spindle speed of 0 rpm. Details of the deformed sheets .......................................................................................... 69 Figure 4.4 FLD with the evolution of strain paths along the final flange for different pre cut hole diameters for a R10 tool radius at 0 rpm .................................................... 72 Figure 4.5 FLD with the evolution of strain paths along the final flange for successful tests with the minimum pre cut hole diameters at 0 rpm ............................................... 73 Figure 4.6 FLD with the evolution of strain paths along the final flange for failed tests xxiv closest to success at 0 rpm ............................................................................................... 74 Figure 4.7 Fractographies of the failure zone for failed tests .......................................... 75 Figure 4.8 Deformation zones on the in-process workpiece ........................................... 76 Figure 4.9 Worst surface quality of the flange obtained in the test with a 10 mm tool radius, 57 mm initial hole diameter and 0 rpm spindle speed: tiny microvoids at the outer surface and grooves in the inner surface follow the forming tool path ................. 77 Figure 4.10 Non-dimensional flange height (h/d f ) of successful hole-flanged sheets for different pre cut diameters at 0 rpm ................................................................................ 79 Figure 4.11 Sheet thickness along the flange formed with spindle speeds of 0 and 1000 rpm ................................................................................................................................... 81 Figure 4.12 Roughness profile Ra of a blank sheet (0° and 90° RD) and deformed flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool radius and 65 mm initial hole diameter .......................................................................... 82 Figure 4.13 Section of specimen R8 with a pre cut hole with a61mm diameter, scanned and compared with the theoretical profile. ........................................................ 83 Figure 4.14 Horizontal and vertical force components vs. tool vertical displacement for successful tests at different spindle speeds. ................................................................ 84 Figure 4.15 Horizontal-to-vertical forces ratio (F h / F z ) for successful and failed tests at different spindle speeds .................................................................................................... 85 Figure 4.16 Vertical Forces for an initial pre cut hole with a 65 mm diameter and punch tool radii of 6, 8 and 10 mm ................................................................................. 86 Figure 4.17 Radial forces for an initial pre cut hole with a 65 mm diameter and punch tool radii of 6, 8 and 10 mm ............................................................................................. 87 Figure 5.1 Evolution of the principal strains along the flange in SPIF and conventional press for a pre cut hole with a diameter of 61 mm and 8 mm tool radius 91 Figure 5.2 Evolution of meridional and circumferential strain along the flange in SPIF and conventional press for a pre cut hole with a diameter of 61 mm and 8 mm tool radius......................................................................................................................... 92 Figure 5.3 Comparison of the mechanisms of failure between both flanging processes. Failed tests with a 60 mm pre cut hole diameter andan 8 mm tool radius by xxxi Chapter 6 presents the main conclusions of this thesis and possible research directions to extend this study in the near future. Figure 0.1 Thesis structure xxxii 1 Analysis of single stage SPIF process applied to the hole flanging operation 1. I NTRODUCTION . S TATE OF THE ART heet metal forming processes have substantial industrial importance in highly competitive sectors that require the fabrication of products with high accuracy and geometrical complexity. Industrial production, particularly in countries with expensive labour, experience a high demand for innovation, flexibility and sustainability in their manufacturing processes to maintain high-quality products at competitive prices. The SPIF process has been proposed as an alternative to the traditional conventional press working process. 1.1. Introduction to conventional flanging by press working When forming a sheet metal part to provide additional support for press fitting or welding to join sheet parts to tubes to improve their appearance or strengthen the edges, the conventional flanging process is one of the last stages of the forming process. This operation is typically employed in the aerospace and automotive industries, in which the flange is generated along the edge of the sheet by the action of a single stroke of a pre-shaped punch (see Figure 1.1). Figure 1.1 Example of flanging by conventional press working (Custompart.net) S 2 State of the art Different flange profiles can be generated by varying the die and punch. Depending on the profile of the edge, the sheet undergoes different strain states during the forming process. Figure 1.2 Types of conventional flanges: A) straight, B) stretch, C) shrink, D) reverse, E) Joggled and F) hole-flanging Figure 1.2 shows different types of flanges. Case A shows a straight flange that is primarily formed under bending around the edge. Case B (referred to as a stretch or concave flange) is formed by a combination of bending and tensile loads. Case C is a convex flange (also referred to as a shrink flange), which is produced by bending and compression loads. Other geometries are usually employed in industry. Cases D and E are a combination of the first three cases. Case F is a particular type of stretch flanging in holes (hole flanging) formed by a combination of circumferential stretching and meridional bending; this type of case is explored in this study. 1.2. Influence of different parameters in conventional hole flanging process In the conventional hole flanging process (Figure 1.3), a clamped blank with a pre-cut hole is plastically deformed and forced to expand. Due to its industrial importance, this process has attracted the interest of researchers who want to understand the deformation process of the flange and the process parameters that control it, such as material properties, anisotropy, punch shape, blank holding force, clearance-thickness ratio or diameter of the initial pre-cut hole. 3 Analysis of single stage SPIF process applied to the hole flanging operation D o D max Punch Die Blank Blank holder Figure 1.3 Schematic of conventional hole flanging, (left) specimen with an initial precut hole and finished hole flanging, and (right) the conventional press working setup. Montanari et al.,2013 These parameters can be classified into two categories: material parameters, i.e., parameters that are directly related to material properties, such as the yield stress, and hardening exponent or anisotropy in the case of metallic materials, and geometrical parameters, which are related to the boundary conditions and loads. The influences of some of these parameters are discussed in the following section. 1.2.1. Material properties Although a wide range of materials exist where this forming process can be analysed, only metallic materials (e.g., as aluminium, steel or titanium alloys) will be considered in this case study. Assuming a generic material hardening law given by =Y+c  , Yamada et al.,1967 investigated the effects of Y and the hardening exponent n in his numeric analysis of the bore expanding test. They noted that circumferential strain increases with n and decreases with Y when the punch load attains its maximum. They also demonstrated that the maximum circumferential stress was located at the edge of the bore, for larger values of n. However, this position can shift to an interior annulus if n decreases and/or Y increases. After ten years, Johnson et al.,1977 experimentally concluded that the process geometry and the plastic 4 State of the art properties of the material influence the instability strain in the hole flanging process and that failure of the flange was due to lip fracture. Some researchers, for instance, Kacem et al.,2015, have investigated the influence of the anisotropy (isotropy, normal, and orthotropic anisotropy) and yield criterion (Von Mises and Hill 1948) on hole flanging of AA1XXX. They concluded that the use of the Hill 1948 yield criterion reasonably explains the analysed geometry and forming parameters. They also evaluated the influence of anisotropy on the total final shape, e.g., the accuracy of the final inner radius and the hole circularity error, the flange thinning and the maximum punch load. 1.2.2. Punch shape Using finite element analysis (FEA), Tang,1981 investigated the flange formation with four different punch shapes (hemispherical, ellipsoidal, nearly flat and conical frustum). He concluded that the strain path at the flange tip during the forming process is independent of the punch shape and the maximum strain at the flange edge for an equal hole size and flange height. The strain path at the flange tip was always proportional and had a strain ratio of -0.5 (pure tensile) (see Figure 1.4 right). However, the maximum load exhibited a high dependence with the punch shape (see Figure 1.4 centre). Figure 1.4 Model (left), punch load vs. displacement (centre) and strain paths at the flange tip for different punch shapes (right). Tang ,1981. Leu et al.,1999 employed an incremental elasto-plastic finite-element method to predict the limiting condition of collar-drawing for three punch shapes (conical, cylindrical and spherical). Table 1-1 presents the experimental results. The authors observed that the critical thickness of the hole for necking for all three punch 5 Analysis of single stage SPIF process applied to the hole flanging operation shapes was equal to the fracture thickness of the material under simple tension. This was used as a necking criterion on the hole periphery for the limiting condition of collar drawing. Table 1-1 Experimental results for a pre cut hole with a diameter of 16 mm for different punch shapes. Leu et al.,1999 Punch type Maximum punch load (kN) Punch stroke (mm) Final diameter of the hole (mm) Thickness on the periphery (mm) Spherical 23.80 28.50 18.29 0.63 Conical 18.31 29.40 18.99 0.65 Cylindrical 41.00 16.50 15.38 0.70 1.2.3. Blank holder effect Krichen et al.,2011 numerically and empirically investigated the blank holding effect on aluminium alloy sheets. Although they concluded that it does not affect the maximum punch load, the shape of the finished product with a residual increase in the outer edge is clearly influenced (see Figure 1.5 right); the maximum load was attained when the test was performed without a blank holder. The author also advised of the use of a blank holding force that at least as great as the reaction force at the beginning of the process. Figure 1.5 Schema of the testing setup (left) and detail of specimens with a blank holding effect (right). Krichen et al.,2011 6 State of the art 1.2.4. Ironing and clearance-thickness ratio Ironing occurs when a flange continuously makes contact with a die and punch during the forming process. Thus, the clearance C between a die and a punch is set lower than a certain value. In the hole flanging practice, one parameter (clearance thickness ratio R c) resumes the three parameters to be considered to undergo a hole flanging process with or without ironing: sheet thickness t, radius of the punch Rp and radius of the die Rd. The clearance C (as shown in Figure 1.6) is defined as the difference between the radius of the die Rd and the radius of the punch Rp. ( =! " −!  ), and the clearance thickness ratio Rc is defined as the ratio between the clearance and the thickness ! % = & ' . Figure 1.6 Schema of the conventional hole flanging process with clearance indication. Kacem et al.,2011 Figure 1.7 FE and experimental shape of flanges for different R c. Kacem et al.,2011 7 Analysis of single stage SPIF process applied to the hole flanging operation Kacem et al.,2011 investigated the critical value of the clearance-thickness ratio in a 1000 series aluminium with a thickness of 2 mm to distinguish between the two hole flanging conditions. Based on this critical value, the authors were able to predict the flange length, thinning of the flange, final shape and punch load. Figure 1.7 shows the flange profile derived from the numerical and experimental campaign that was performed for different Rc values.Figure 1.8shows the flange length and thinning evolutions according to the evaluated Rc. Figure 1.8 Flange length and thinning evolution for different R c. Kacem et al.,2011 Figure 1.9 Distribution of the stress triaxiality in the conventional hole flanging process on aluminium A1050-H14 with and without ironing. Kacem et al.,2013 8 State of the art In a subsequent study, Kacem et al.,2013 investigated the prediction of damage on aluminium A1050-H14 and A6061-O with different parameters, such as the clearance-thickness ratio (Rc) or ironing. Based on the results of an experimental and numerical analysis, they revealed that ironing decreases the damage to the flanged parts since it produces a compressive stress state that tends to close existing voids whereas the clearance-thickness ratio and the diameter of the initial hole significantly affect the damage state of the flange. Figure 1.9 shows the triaxiality distribution in two hole flanging conditions: without ironing (upper) and with ironing (lower). Due to the imposed geometry between the punch and the die in the resulting flange for low values of the clearance thickness ratio (Rc), a compressive strain state is induced in the flange, which prevents void growth and coalescence. 1.2.5. Initial pre cut hole Huang et al.,2001 numerically investigated the initial pre cut hole effect on the lip wall thinning of a steel sheet. Changing the initial pre cut hole diameter for the same cylindrical punch and monitoring the lip thinning, the authors realized that the maximum reduction in the wall thickness occurred along the periphery of the expanded hole and the tensile stress in the circumferential direction at this location was responsible for failure by cracking or tearing. Thus, they established a useful tool design for engineers by plotting the relationship between the initial hole diameters and expanded the hole diameters for fractured and non fractured specimens. Figure 1.10 shows this relationship by a straight line. Figure 1.10 Failure map for the conventional hole flanging process at a fixed punch diameter. Huang et al.,2001 15 Analysis of single stage SPIF process applied to the hole flanging operation SPIF, which is the simplest variant of ISF, has important drawbacks regarding the dimensional accuracy compared with TPIF. Its use is not in the scope of industry; thus, few patents have been registered in recent years about the process to increase the dimensional accuracy and reduce the springback effect of the blanks. For the patent requested in 2005 by the Univ. of Leuven, Belgium, [EP 1899089], as shown in Figure 1.19, the SPIF device is equipped with a heating device. Figure 1.19 Patent number [EP 1899089] from Univ. Leuven 2005 Although car manufacturers initially boosted the study of this process by registering a large amount of patents and manufacturing a variety of pieces (reflective surface of prototype headlights; heat/noise shield, which is used over exhaust manifolds (Figure 1.20, Jeswiet et al.,2001)); service panels (Figure 1.21, Amino et al.,2002, 2014); hood and fender (Rodriguez et al.,2006), which is used by other industries;customized ankle support (Ambrogio et al.,2005) or cranial plate (Figure 1.22, Bagudanch et al.,2015) in biomedical applications and appliances (solar cooker, Jeswiet et al.,2005) and for processing recycling panels, as demonstrated by Jackson et al.,2008 and Takano et al.,2008. 16 State of the art Figure 1.20 Formed by SPIF headlights and manifolds by Jeswiet et al.,2001 Figure 1.21 Service panels before and after SPIF forming by Amino et al.,2002, 2014 Figure 1.22 Cranial plate capture (left) and SPIF cranial plate (right) by Bagudanch et al.,2015 The following advantages make this process variant useful for prototyping in small- and medium-sized production, apart from an increase in formability: lower cost than the conventional sheet metal forming process, can be executed in a conventional CNC machine, flexible design, changes can be easily implemented, production directly from CAD file and small force requirements compared with conventional sheet metal forming process in the case of small thickness metal sheets or soft materials. Although most research groups employ an adapted CNC machine, probably for economic reasons, some research groups have designed new devices that can support 13kN axial force and 6.5 kN lateral force with 6 17 Analysis of single stage SPIF process applied to the hole flanging operation d.o.f., as shown in Figure 1.23, due to the development of harder materials and the need to test metal sheets with higher thicknesses. Figure 1.23 CAD model (left) and built machine (right) of SPIF-A. Alves de Sousa et al.,2014 1.5. Formability in SPIF Formability in sheet metal forming is usually evaluated using a Forming Limit Diagram (FLD). A FLD represents the in-plane major vs. minor strain of a sheet and can include different curves to divide the strain space, such as the Forming Limit Curve (FLC) for the onset of necking and the Fracture Forming Limit (FFL) for fracture. The FFL shape is a straight line that corresponds to a fracture failure mode by tension (Mode I), which is typical of metal sheets. Recently, Isik et al.,2014 experimentally determined by SPIF, torsion and in-plane shear tests of AA1050-H111 sheets a third failure curve for fracture by in-plane shear (Mode II). This curve is known as the Shear-FFL (SFFL). Figure 1.24 shows a schema of experiments and experimental results in the FLD including FFL and SFFL. Using a conical or pyramidal geometry, the authors showed the enhanced formability limit of SPIF processes above the FLC. Many factors involved in the forming process can interfere in the enhancement of the formability of this process. 18 State of the art Figure 1.24 Geometries (left) of tests used to calculate (right) SFFL and FFL in AA1051-H111, isik et al.,2014 1.5.1. Suppression of necking To explain this enhanced formability, numerous papers discuss different mechanisms that affect, e.g., localized bending, via thickness shear and triaxiality, such as Martins et al.,2008, Emmens et al.,2009a, 2009b and Malhotra et al.,2012. Many authors have investigated the factors that produce these beneficial effects. Silva et al.,2011 based their research on the tool radius effects and the stabilizing effect of dynamic bending under tension on AA1050-H111. With a large tool radius, this process seems to be capable of increasing the forming limit curve above that commonly found in stamping to ensure localization by necking. In the case of a small tool radius, the stabilizing effect is not sufficient for ensuring localization and a failure mechanism change to fracture without necking. In their research on AISI304, Centeno et al.,2013 highlighted that the postponed or suppressed necking in ductile metal sheets deformed by SPIF, which enable failure directly by ductile fracture, is caused by not only the tool radius bending effect but also an extensive range of process parameters (tool diameters, step downs, and friction conditions). Fang et al.,2014 proposed an analytical model of the deformation and fracture of the sheet in AA1100that considers the bending and strain hardening effect. 19 Analysis of single stage SPIF process applied to the hole flanging operation Figure 1.25 Fracture localization due to the bending effect in AA1100. Fang et al.,2014 Fang theoretically proves that fracture occurs first on the outer side of the wall at the transition area between the contact and the non-contact zone due to the bending effect (See Figure 1.25). The strain hardening has a positive or negative effect depending on the thickness ratio. Li et al.,2015 numerically analysed the deformation mechanics and the forming force in SPIF of AA7075-O sheets and concludes that the SPIF deformation process is a combination of stretching, bending and shearing. This combination oscillates depending on the process variables, such as wall angle, friction conditions, and tool diameter. Figure 1.26 Percentage contribution of each deformation mode for four elements under varying wall angles. Maqbool et al.,2017 20 State of the art Figure 1.27 Percentage contribution of each deformation mode for two elements under the influence of the a) tool diameter, b) thickness, c) friction and d) step down. Maqbool et al.,2017 As indicated by Maqbool et al.,2017 in their study about plastic energy dissipation using an analytical approach, the dominant deformation mode depends on the process variables. According to the desired target, they can be employed in the design forming strategy. Figure 1.27 and Figure 1.26 show the percentage contribution of each deformation mode (shear, bending or membrane) to different process variables, such as tool diameter, thickness, friction, step down and wall angle. Some mechanisms depend on factors, such as mechanical properties or thickness, whereas other mechanisms are related to factors that are based on the interaction of the tool with the blank, such as tool shape, spindle speed or lubrication conditions. Other mechanisms depend on the tool trajectories: step down, feed rate or tool strategies. These factors are extensively analysed in the next three sectionsInfluence of material properties 1.5.2. Influence of material properties  Material properties. A relationship between material properties and formability is observed in conventional press working. Fratini et al.,2004 investigated the influence of six mechanical parameters on the enhanced formability of the SPIF process for six materials: K, strength coefficient; n, strain hardening coefficient; Rn, normal 21 Analysis of single stage SPIF process applied to the hole flanging operation anisotropy index; UTS, ultimate tensile strength and A%, percentage elongation. Table 1-2 (left) shows the list of evaluated materials and properties. To study the enhanced formability, Fratini needed a reference value of formability to compare the FLD0 calculated according to Geiger et al.,2003 (FLC under plane strain conditions) with the FLD0incre (FFL under plane strain conditions) obtained from tests of truncated cone or pyramids, as these SPIF strain paths are typically very close to the major strain axis in a forming limit diagram. Table 1-2 Properties of materials investigated (left) and influence of the inputs variables and interactions on FLD 0 and FLD 0incre (right). Fratini et al.,2004 According to the results of their investigation, as presented in Table 1-2 (right), Fratini´s model confirmed that the most important material parameters that influence conventional press working or FLD0 were the hardening coefficient and normal anisotropy, whereas the highest influence on material formability for incremental forming processes is derived from the hardening coefficient. Also the strength coefficient and percentage elongation have a relevant influence. Since process mechanics in incremental forming is characterized by stretching and local thinning, the higher is the strain hardening coefficient, the larger is the material formability, which is consistent with the fact that the strain hardening coefficient indicates an indirect measurement of material capability to undergo thinning without plastic collapse. 22 State of the art  Material thickness Hussain et al.,2007 proposed a method that was ratified one year later by Silva et al.,2008, to test the thinning limits of sheet metals by performing a series of tests under Negative Incremental Forming conditions (also referred to as Dieless Forming) and comparing these values with values predicted with the Law of Cosines, as presented in Figure 1.28. Truncated cone or pyramids of fixed or variable wall angles are typical selected geometries. Figure 1.28 Schematic of the Law of Cosines. Hussain et al.,2007 Material thickness is present in the Law of Cosines (= 3 cos7) and was considered in the experimental investigation performed in 2013 by Liu et al. for three thicknesses of AA7075-O aluminium sheets: 1.02 mm, 1.6 mm and 2.54 mm. They compared these three thicknesses from several points of view to investigate the influence of mechanical properties, formability and forces. They concluded that the material thickness clearly influences the formability of the SPIF process if the phenomenon of anisotropy increases with the material thickness. In recent investigations, Hussain et al.,2013 tested variable wall angle truncated cones in anextensive range of AA2024-O thicknesses, which enhances the importance of choosing an accurate tool size for the proper thickness; thus, the tool size plays a role. 1.5.3. Influence of the forming tool SPIF processes are characterized by their low production rates as a result of the local incremental nature. As a consequence of this local forming process, the 23 Analysis of single stage SPIF process applied to the hole flanging operation surface finish is penalized and the formability is highly dependent of the friction conditions. Thus, the tool shape and size, spindle speed, and lubrication conditions are strongly related.  Tool shape The simplest tool shape in SPIF is the hemispherical end tool. In 2013, Cawley et al. compared angle and parabolic end tools with flat and hemispherical end tools geometries (see Figure 1.29) with the objective of improving the forming limits and expanding the range of potential applications. The test was performed on a variable wall angle conical specimen with a 1.59 mm thick Al-3003-O. Figure 1.29 Angle, flat, parabolic and hemispherical tool shapes tested by Cawley et al.,2013 In their research, they reaffirmed that a reduced radius contact surface increased the forming limit and worsened the surface quality. They concluded that the use of an end tool or a combination of end tools is the desired option. Figure 1.30 a) Rigid hemispherical tool, b) vertical roller ball and c) oblique roller ball. Lu et al.,2014 24 State of the art To study the influence of friction on the surface finish, Lu et al.,2014 experimentally and theoretically investigated the differences in surface finish between two types of spherical forming tool: a rigid hemispherical tool and a roller ball in two geometries (vertical and oblique), as shown in Figure 1.30. Although Lu et al.,2014 investigated four grades of aluminium sheet materials, including AA2024, AA5052 and AA6111, AA1100 was employed to perform a series of tests to evaluate the surface roughness Rz and Ra in truncate cone geometries, as shown in Figure 1.31. Figure 1.31 a) Truncate cone geometry of surface roughness tests, b) part processed with a rigid tool, c) part processed with VRB and d) with ORB. Lu et al.,2014 Figure 1.32 Result of surface roughness tests in terms of Rz and Ra. Lu et al.,2014 31 Analysis of single stage SPIF process applied to the hole flanging operation As shown in Figure 1.41, Mirnia et al.,2014 obtained an improved thickness distribution of a 70o truncated cone in AA1050 by sequential limit analysis compared with single-stage or conventional multi-stage SPIF, which decreases the forming time. Shamsari et al., 2018 proposed a less sophisticated two-stage hybrid deformation strategy. Compared with SPIF in a single stage, they demonstrated an increase in depth and a decrease in thinning in a truncated cone by the SPIF process, which induces an equibiaxial strain state by hydraulic bulging. In recent years, due to numerical simulations and the use of five-axis CNC machines, a large number of authors investigated tool path strategies to increase the dimensional accuracy, such as Giraud-Moreau et al.,2018 and Fiorentino et al.,2018, and optimize the dimensional accuracy by an iterative correction of the tool path. The thickness distribution and process limits have also been investigated with these new machines and methodologies. 1.6. Hole-flanging by SPIF in a single stage Although the majority of research has been based on a truncated cone, pyramid or variable wall angle conical frustum, current studies show the capability of SPIF to successfully perform circular hole flanging, where a sheet with a pre cut hole is deformed by a forming tool that progressively produces a smooth round flange following a pre-established trajectory,. The pioneering research on hole flanging by SPIF is relatively recent, as presented by Cui and Gao in 2010. In their research, they perform successful hole flangings using different multi-stage strategies compared with the thickness distribution obtained in each case, as shown in Figure 1.42. Cui and Gao analysed the formability of the process in terms of the LFR, which demonstrates that the use of strategy “a” is more versatile than the other two strategies and achieves a larger LFR and more uniform thickness distribution. Two years after Cui’s study, Centeno et al.,2012 attempted to answer some questions regarding the formability aspects that Cui and Gao did not examine, in particular, the influence of a pre cut hole in the formability of SPIF. They concluded that small pre cut holes did not have any influence on the formability of the process if the plastic deformation does not reach the vicinity of the hole edge. However, a new mode of failure is triggered in the contrary case, which entails a 32 State of the art combination of in-plane stretching in plane strain conditions with bending around the perimeter of the flange, which causes a suppression of necking prior to failure (see Figure 1.43 a). Figure 1.42 Tool path strategies and corresponding thickness distributions for an initial pre cut hole with d o =35.5 mm and final hole diameter of d f =65 mm in AA1060. Cui et al.,2010 Figure 1.43 a)Experimental strains in conical hole flanging with multi-stage strategies in AA1050 b) and selected steps in the successful production of cylindrical hole flanging. Centeno et al., 2012 33 Analysis of single stage SPIF process applied to the hole flanging operation Figure 1.43 b) also depicted the strain distribution of outer surface points along a meridional cut in the FLD. The meridional strain distribution follows the plane strain behaviour at the bottom of the specimen due to the equibiaxial strain state near the hole edge. Due to the particular geometry of the holes, most studies on hole flanging have been devoted to analysing the sheet formability via the LFR, and few studies have applied the Forming Limit Curve (FLC) as an analysis tool, such as studies by Martins and co-workers (Silva et al.,2013, Montanari et al.,2013 and Cristino et al.,2015). Although the FLC is currently the most useful tool for evaluating the formability of metal sheets during the design process of press working operations, these studies have not provided a detailed description of the flange deformation process or a comprehensive study of the influence of the punch profile. Figure 1.44 depicts some results of the strain analysis in hole flanging by SPIF and conventional FLD. Figure 1.44 Strain distributions along meridional directions of hole-flanged parts in a) AA1050-H111 and b)titanium (grade 2) under press working conditions and multi-stage SPIF. Silva et al.,2013 Despite their inherent advantages in terms of simplicity, flexibility and sustainability, multi-stage hole flanging processes by SPIF have a large deficit in terms of time consumption. In this regard, some attempts have been taken reduce production times by performing hole flanging processes by high-speed SPIF. For instance, in a study by Bambach et al., 2014, the process was accelerated using a rotating tool set, on which the forming tool was radially and axially displaced, as shown in Figure 1.45. 34 State of the art Figure 1.45 High-speed incremental hole flanging: tool designs a), setup b), parameters c) and specimens before and after performing. Bambach et al.2014 Others authors have decided to create new featured tool and tool path strategies to accelerate the process. In Figure 1.46, a basic tool trajectory is compared with the new one tool trajectory and the forming forces and timing. Figure 1.46 Different tool path strategies, forming forces and timing for both tools. Cao et al., 2016 Apart from the studies by Borrego et al.,2015, 2016, a systematic study about the capability of performing hole flanging by SPIF in a single stage and an analysis of the mechanics of the process in terms of deformations induced in the part or about the failure modes that define its formability are lacking. 35 Analysis of single stage SPIF process applied to the hole flanging operation 2. M ATERIAL CHARACTERIZATION he aluminium alloy sheet AA7075-O with a thickness of 1.6 mm, which is typically employed in aeronautical applications, is the material selected for this study. The material characterization includes mechanical parameters, such as the flow curve, Young’s modulus, ultimate tensile strength, Lankford coefficients and Forming Limit Curve, which is an extensively applied tool in mechanical forming for evaluating formability. 2.1. Mechanical parameters Generally, the first step after receiving a batch of new material is to perform a tensile test to determine the mechanical properties and compare them with the properties of other batches. A data sheet provides information about the chemical composition and basic mechanical properties but does not provide the flow curve of the desired strain levels. 2.1.1. Tensile test A tensile test is the most common test for evaluating the elastic-plastic behaviour of a material and is usually performed to calculate Young´s modulus, yield stress, ultimate tensile stress and Lankford coefficients for three different directions (0º rolling, 90º transverse and 45º diagonal direction). A series of tensile tests was carried out in a universal testing machine at room temperature. The gauge length and width of the tensile specimens were 60 and 12.5 mm respectively according to the standards ASTM E8M-08 (see specimen dimension in Figure 2.1). The mechanical properties at 0º, 45º and 90º with respect to the rolling direction were also obtained. An extensometer of 25 mm longitudinal reference length and +/-5 mm range was used. Three tests for each direction were performed. T 36 Material characterization Figure 2.1 a) Geometry of the tensile test specimen and b) strain-stress curves in three different directions Figure 2.1depicts the true stress-strain curves at 0º, 45º and 90º. The mean values obtained for the yield stress (Y), ultimate tensile strength (UTS), and Young's modulus (E) are displayed in Table 2-1. An average value in the sheet plane is also presented as a reference. Table 2-1 Material properties of the AA7075-O Sheet direction Y (MPa) UTS (MPa) E (GPa) Rolling (0º) 110 214 63.7 Diagonal.,45º) 108 212 61.6 Transverse (90º) 112 218 64 Average: . 8 9  . :; 9 . <8  109.5 214 62.5 (b) (a) 37 Analysis of single stage SPIF process applied to the hole flanging operation According to Gilmour et al., 2004, the most accurate correlations between the numerical predictions and the experimental tests for the forming processes that involve high strain levels are achieved by evaluating the Lankford coefficients using plastic strains at levels near the end of uniform strain behaviour. The Lankford coefficients were obtained according to the ASTM 517-00 standard at an engineering strain level of 0.11, following equation (2-1). Once applied, the volume constancy assumes the following form:  =   =   = ≅ −   =   = +   = (2-1) where    and    are the longitudinal principal plastic strain and transverse principal plastic strain, respectively, in the tensile specimen, and    is the normalor thickness) strain plastic evaluated by volume constancy from the in plane strains;   = - (   +   ). Table 2-2 lists the mean values from three tests of the Lankford coefficients at 0º, 45º and 90º and the conventional normal anisotropy coefficient. These values are consistent with the experimental findings of Liu et al., 2013 and Durante et al., 2009 for the same type of aluminium alloy. Table 2-2 Lankford coefficients for 0°, 45° and 90° 2.1.2. Bulge test According to the tensile test results, only strain levels of approximately 0.12 were achieved. A series of bulge test was performed to extrapolate the flow curve from 0.12 to higher strain levels. A bulge test consists of an induced equal biaxial strain state by inflating a sheet plate under hydraulic pressure through a circular die (see Figure 2.2). A complete setup that comprises an aluminium sheet plate to be tested between a backing plate and an annular blank holder or die was developed. The backing plate was provided for a container of hydraulic fluid. A blank holder is screwed to the backing plate,   °   °   °   ° + 2   ° +   ° 4 0.65 0.95 0.81 0.85 38 Material characterization and a draw bed is machined between both the blank holder and the draw bed to provide sufficient clamping force. Figure 2.2 Section of the bulge test setup (recreation) The standard ISO 16808:2014 describes the methodology for calculating the strains and curvature radius at the top of the dome when using video correlations systems (recreation shown in Figure 2.3). Once the test is computed and the top of the dome is localized at the last image before breaking, two circular areas of radius   and   are defined around the top:   = (0.125±0.025)*d (2-2)   = (0.05±0.01)*d where d is the die diameter, which is 100 mm in this study. r1 is used to calculate the area where to perform the best fitting sphere to get the radius of curvature and the strains   =  at the dome of the specimen are estimated using a circular surface of radius r2 Pressure measurements were directly obtained from a transducer at the inlet of the container. According to Lazarescu et al., 2011, the instantaneous average thickness at the top of the dome can be estimated by assuming material incompressibility, an isotropic behaviour and a spherical shape of the dome using:  =  3 exp ( − 2   ) (2-3) Thus, the equivalent stress and strains can be calculated by   = =E  ' ;   = ln ( ' '  ) . (Lazarescu et al.,2011) (2-4) 39 Analysis of single stage SPIF process applied to the hole flanging operation Figure 2.3 Configuration of cameras for the bulge test under the video correlation system Figure 2.4 shows the true stress-strain curve obtained in the bulge test. The true stress-strain curves at 0º in the tensile test is also depicted for comparison. The bulge curve successfully reproduces the tensile behaviour and is able to extend the stress-strain curve to 0.7 of the equivalent strain. Figure 2.4 Equivalent stress vs equivalent strain for tensile and bulge tests Material characterization 2.1.3. Incremental forming is an extraordinary demanding process in terms perform accurate numerical simulations, which is crucial for obtaining flow curves with high strain levels. The stack compression test is an easy method for obtaining flow curves with higher strain levels due to the absence of necking. In th is test, a pile of discs was axially compressed by two polished and lubricated parallel plates under the action of a hydraulic cylinder ( compression force and vertical displacement were monitored during the test. The test follo the aspect ratio of the stack, which was suggested in are the height of the stack and the diameter of the stack, respectively. In this study, according to the conclu of sheets or plates can be successfully utilized, but due to the difficulties of stacking more than five discs in a stable manner, a pile of five discs with a thickness of 1.6 mm and a diameter of 2.6 shows the agreement between the bulge tests and the stack compression test, which enables the material to be characterized to a 1 strain level. Material characterization 2.1.3. Stack compression test Incremental forming is an extraordinary demanding process in terms perform accurate numerical simulations, which is crucial for obtaining flow curves with high strain levels. The stack compression test is an easy method for obtaining flow curves with higher strain levels due to the absence of necking. is test, a pile of discs was axially compressed by two polished and lubricated parallel plates under the action of a hydraulic cylinder ( compression force and vertical displacement were monitored during the test. Figure The test follo wed the recommendation of the ASTM E9 the aspect ratio of the stack, which was suggested in are the height of the stack and the diameter of the stack, respectively. In this study, according to the conclu sions of Alves of sheets or plates can be successfully utilized, but due to the difficulties of stacking more than five discs in a stable manner, a pile of five discs with a thickness of 1.6 mm and a diameter of shows the agreement between the bulge tests and the stack compression test, which enables the material to be characterized to a 1 strain level. Stack compression test Incremental forming is an extraordinary demanding process in terms perform accurate numerical simulations, which is crucial for obtaining flow curves with high strain levels. The stack compression test is an easy method for obtaining flow curves with higher strain levels due to the absence of necking. is test, a pile of discs was axially compressed by two polished and lubricated parallel plates under the action of a hydraulic cylinder ( compression force and vertical displacement were monitored during the test. Figure 2.5 Stack compression test setup wed the recommendation of the ASTM E9 the aspect ratio of the stack, which was suggested in are the height of the stack and the diameter of the stack, respectively. In this study, sions of Alves et al., of sheets or plates can be successfully utilized, but due to the difficulties of stacking more than five discs in a stable manner, a pile of five discs with a thickness of 1.6 mm and a diameter of 15 mm ( shows the agreement between the bulge tests and the stack compression test, which enables the material to be characterized to a 1 strain level. Stack compression test Incremental forming is an extraordinary demanding process in terms perform accurate numerical simulations, which is crucial for obtaining flow curves with high strain levels. The stack compression test is an easy method for obtaining flow curves with higher strain levels due to the absence of necking. is test, a pile of discs was axially compressed by two polished and lubricated parallel plates under the action of a hydraulic cylinder ( compression force and vertical displacement were monitored during the test. Stack compression test setup wed the recommendation of the ASTM E9 the aspect ratio of the stack, which was suggested in 1 ≤ are the height of the stack and the diameter of the stack, respectively. In this study, et al., 2011, raw materials supplied in form of sheets or plates can be successfully utilized, but due to the difficulties of stacking more than five discs in a stable manner, a pile of five discs with a 15 mm ( h o/do ≈0.5) was employed. shows the agreement between the bulge tests and the stack compression test, which enables the material to be characterized to a 1 strain level. Incremental forming is an extraordinary demanding process in terms of strains to perform accurate numerical simulations, which is crucial for obtaining flow curves with high strain levels. The stack compression test is an easy method for obtaining flow curves with higher strain levels due to the absence of necking. is test, a pile of discs was axially compressed by two polished and lubricated parallel plates under the action of a hydraulic cylinder ( Figure 2. 5 compression force and vertical displacement were monitored during the test. Stack compression test setup wed the recommendation of the ASTM E9 - 09, with the exceptionof ≤ ho/do ≤3, where are the height of the stack and the diameter of the stack, respectively. In this study, 2011, raw materials supplied in form of sheets or plates can be successfully utilized, but due to the difficulties of stacking more than five discs in a stable manner, a pile of five discs with a ≈0.5) was employed. shows the agreement between the bulge tests and the stack compression test, which enables the material to be characterized to a 1 strain level. 40 of strains to perform accurate numerical simulations, which is crucial for obtaining flow curves with high strain levels. The stack compression test is an easy method for obtaining is test, a pile of discs was axially compressed by two polished and lubricated 5 ). Both the compression force and vertical displacement were monitored during the test. 09, with the exceptionof ≤3, where ho and do are the height of the stack and the diameter of the stack, respectively. In this study, 2011, raw materials supplied in form of sheets or plates can be successfully utilized, but due to the difficulties of stacking more than five discs in a stable manner, a pile of five discs with a ≈0.5) was employed. Figure shows the agreement between the bulge tests and the stack compression test, 47 Analysis of single stage SPIF process applied to the hole flanging operation 3. H OLE FLANGING BY CONVENTIONAL PRESS WORKING he aim of this chapter is to contribute to the understanding of the factors that control the formability of the sheet during hole flanging operations using conventional press working. The deformation process of the flange and the influence of the initial hole size and punch profile in the failure and final geometry of the flange are discussed. A series of tests using AA7075-O metal sheets was performed and analysed using three different cylindrical punches with different profile radii (10 mm, 8 mm and 6 mm) and a hemispherical punch (R=47.9 mm). The evolutions of the principal strains along the flange were measured using an optical 3D forming analysis system (ARGUS®). The formability of the flange was analysed via the traditional LFR and conventional FLC. The findings indicated that the FLC was not suitable for analysing the formability of the material along the flange wall, with the exception of the tip of the flange. The real influence of bending induced by the punch and the apparent independence of the punch profile on the flange formability was elucidated. 3.1. Experimental procedure The tests were performed using an Erichsen 142-20 machine with a blank holder and backing plate with a diameter of 100 mm inner. Figure 3.1 schematically depicts the hole flanging process performed in the laboratory. Figure 3.1 Conventional hole flanging process scheme T 48 Hole flanging by conventional press working As shown in the scheme, during the flanging process, the punch expands the material around the pre cut hole in the blank to form a hollow flange. The operation attempts to achieve the final flange with a single pass of the forming tool and is successful when a complete flange that is free of defects is obtained. Table 3-1 shows the proposed experimental campaign. The blanks were circular, with an external diameter of 180 mm and central pre cut holes with different diameters do. The central holes were milled and subsequently grinded with fine grit sandpaper to eliminate any burrs. Table 3-1 Experimental campaign for hole flanging by conventional press working Figure 3.2 Set of punches with different edge radii for hole flanging tests by conventional press working (6 mm, 8 mm, 10 mm and 47.9 mm) Pre cut diameter hole, do (mm) Die diameter Df (mm) Punch diameter, Dp (mm) Edge radius R (mm) 35 100 95.8 41 58 6 59 8 60 10 61 47.9 65 72 83 49 Analysis of single stage SPIF process applied to the hole flanging operation A series of punches with an equal circumferential diameter Dp and different edge radii R was used to analyse the effect of the punch profile on the formability. A circumferential diameter of 95.8 mm and four edge radii of 6 mm, 8 mm, 10 mm and 47.9 mm (hemispherical punch) were selected (see Figure 3.2). The gap between the die and the punch was slightly larger than the sheet thickness to avoid ironing during the forming process. The punches, die and blank holder were fabricated from UNE F1430 steel, tempered and hardened to 60 HRC. The punch velocity was set to 1 mm/s. The contact between the blank and the punch was lubricated with Vaseline. The blank holding force was set to 85 kN, which was sufficiently high to avoid the blank from drawing in. According to Krichen et al., 2011, the blank holding conditions do not influence the maximum punch load. However, an insufficient holding force may modify the finished flange due to springback effects. Therefore, the use of holding forces that are higher than the maximum reaction force experienced by the punch during the test is always advised. To study the flange deformation and failure mechanisms, the strains at the outer sheet surface were measured using a circle grid analysis. Blanks were electro etched with a grid of dots with a diameter of 1 mm and centres spaced by 2 mm and the optical 3D forming analysis system ARGUS®, which is used to automatically compute the principal strains at the sheet surface. Measurements were performed at the end of the test or, in the case of failure, after the specimen failed. The final flange height was also measured to evaluate the geometric capabilities of the process. 3.2. Specimen marking for strain measurement The marking process and limitations in the strain measurement with ARGUS® are discussed in this section. 3.2.1. Electrolytic marking process The electro etching machine used in this study was an EU-Classic 300 from Ötsling Marking Systems, which provides AC/DC current from 0-24 V. The current flow and voltage selected combined with the electrolyte and felt of the marking head leads to a marking pattern based on black or white dots on the same AA7075-O. Hole flanging by conventional press In our study, a stencil of a dot g was employed. A dot grid with a diameter of 0.5 mm and spacing of 1 mm was checked for better accuracy. Although this stencil is indicated to strain gradient areas, it easily disappears when a high strain lev According to the manufacturer, the selected combination of electrolytic, felt, current flow and different voltages for black markings did not meet our expectations for high process. They d points and the shiny aluminium sheet background. The solution passed by finding a method to increase the contrast by increasing the black colour of the points or whitening the background of To increase the black colour of the points, a sand blasting process of the surface prior to electro etching increased the contrast; however, a compression strain state was induced in the sand Figure flanging by conventional press Figure In our study, a stencil of a dot g was employed. A dot grid with a diameter of 0.5 mm and spacing of 1 mm was checked for better accuracy. Although this stencil is indicated to strain gradient areas, it easily disappears when a high strain lev According to the manufacturer, the selected combination of electrolytic, felt, current flow and different voltages for black markings did not meet our expectations for high - demanding strain states in hole flanging by the SPIF process. They d id not provide sufficient contrast between the black colour of the points and the shiny aluminium sheet background. The solution passed by finding a method to increase the contrast by increasing the black colour of the points or whitening the background of To increase the black colour of the points, a sand blasting process of the surface prior to electro etching increased the contrast; however, a compression strain state was induced in the sand - blasted surface. This solution was discar Figure 3.4 Examples of black on white and white on black markings flanging by conventional press working Figure 3.3 Electro etching setup In our study, a stencil of a dot g rid with a diameter of 1 mm and spacing of 2 mm was employed. A dot grid with a diameter of 0.5 mm and spacing of 1 mm was checked for better accuracy. Although this stencil is indicated to strain gradient areas, it easily disappears when a high strain lev According to the manufacturer, the selected combination of electrolytic, felt, current flow and different voltages for black markings did not meet our demanding strain states in hole flanging by the SPIF id not provide sufficient contrast between the black colour of the points and the shiny aluminium sheet background. The solution passed by finding a method to increase the contrast by increasing the black colour of the points or whitening the background of the aluminium sheet. To increase the black colour of the points, a sand blasting process of the surface prior to electro etching increased the contrast; however, a compression strain state blasted surface. This solution was discar Examples of black on white and white on black markings working Electro etching setup rid with a diameter of 1 mm and spacing of 2 mm was employed. A dot grid with a diameter of 0.5 mm and spacing of 1 mm was checked for better accuracy. Although this stencil is indicated to strain gradient areas, it easily disappears when a high strain lev el is reached. According to the manufacturer, the selected combination of electrolytic, felt, current flow and different voltages for black markings did not meet our demanding strain states in hole flanging by the SPIF id not provide sufficient contrast between the black colour of the points and the shiny aluminium sheet background. The solution passed by finding a method to increase the contrast by increasing the black colour of the points or the aluminium sheet. To increase the black colour of the points, a sand blasting process of the surface prior to electro etching increased the contrast; however, a compression strain state blasted surface. This solution was discar Examples of black on white and white on black markings Electro etching setup rid with a diameter of 1 mm and spacing of 2 mm was employed. A dot grid with a diameter of 0.5 mm and spacing of 1 mm was checked for better accuracy. Although this stencil is indicated to strain gradient el is reached. According to the manufacturer, the selected combination of electrolytic, felt, current flow and different voltages for black markings did not meet our demanding strain states in hole flanging by the SPIF id not provide sufficient contrast between the black colour of the points and the shiny aluminium sheet background. The solution passed by finding a method to increase the contrast by increasing the black colour of the points or To increase the black colour of the points, a sand blasting process of the surface prior to electro etching increased the contrast; however, a compression strain state blasted surface. This solution was discar ded. Examples of black on white and white on black markings 50 rid with a diameter of 1 mm and spacing of 2 mm was employed. A dot grid with a diameter of 0.5 mm and spacing of 1 mm was checked for better accuracy. Although this stencil is indicated to strain gradient According to the manufacturer, the selected combination of electrolytic, felt, current flow and different voltages for black markings did not meet our demanding strain states in hole flanging by the SPIF id not provide sufficient contrast between the black colour of the points and the shiny aluminium sheet background. The solution passed by finding a method to increase the contrast by increasing the black colour of the points or To increase the black colour of the points, a sand blasting process of the surface prior to electro etching increased the contrast; however, a compression strain state Examples of black on white and white on black markings 51 Analysis of single stage SPIF process applied to the hole flanging operation For whitening the background prior to electro etching, a Keller solution was applied (93.5% H2O, 1.5% HCl, 2.5% HNO3 and 2.5% HFl). After plunging the sheet blank in the solution and electro etching, perfect black dots were visible in the white background (see Figure 3.4 left). This technique yields positive results in terms of sufficient contrast between the black and white colours for high strain levels and will be employed in the hole flanging by SPIF discussed in section 4. In the case of a conventional press forming test, where lower strain levels were achieved at the outer surface of the specimen, both black on white or white on black marking can be applied. In our case, a white on black marking was performed using deep marking settings. Note that ARGUS® reads the shiny aluminium patina as a black background when no direct light is falling on it. The optical 3D forming analysis system ARGUS® computes the strains after the hole flanging process by comparing the centre distance of the previously electro etched dotted pattern in the outer surface of the specimen with the centre distance of the initial dotted pattern, which is an internal parameter of the software that can be adjusted according to the dot grid stencil used during the marking process. (ARGUS User Manual v6.2 rev-a, 17/09/2009.) Figure 3.5 Specimen before and after ARGUS computation; an example in conventional and SPIF processes Figure 3.5 shows a typical setup for measuring the strains using the ARGUS® system and the major principal strain contour map. Specimens for conventional and incremental hole flanging are shown. The main drawback of the software is that it cannot compute strains near the edges of the sheet due to the lack of data for interpolation. Thus, in this study, only the strains at the hole edge have been calculated from direct measurement of the initial and final perimeters of the hole and the initial and final sheet thickness and assuming material volume constancy. 52 Hole flanging by conventional press working 3.3. Experimental results In this section, the tests results are described. The influence of the punch profile on the deformation of the flange is discussed. Table 3-2 Series of conducted conventional press working hole flanging tests Edge radius, R (mm) Initial pre cut hole diameter, do(mm) 35 41 58 59 60 61 65 72 83 6 F F F/N N O O O O 8 F F F/N N O O O O 10 F F/N N O O O O 47.9 F F/N N O O O O Table 3-2 summarizes the results of the hole flanging process using AA7075-O produced by conventional press working in blanks with different pre cut hole diameters do.The experimental plan is shown in Table 3-1. The different punches are identified by their edge radius R. The label “O” indicates a successful test in terms of obtaining a complete hole-flanged part that was free of defects; “N” indicates tests that exhibited necking without fracturing the flanges; and “F” indicates specimens with fractured flanges. Figure 3.6 Flanges from pre cut hole diameters of 59 (fractured), 60 (necked) and 61 mm (successful) with a 6 mm edge radius 53 Analysis of single stage SPIF process applied to the hole flanging operation A series of increasing values of do was tested to determine the successful flanges. Intermediate pre cut holes were tested until a difference equal or less than 1 mm in diameter was reached between the successful tests and the failed tests. The necking situations were verified using at least two different tests. Figure 3.6 shows tests for the pre cut holes with a 59 mm, 60 mm and 61 mm diameter using a punch with a 6 mm edge radius. A successful flange was obtained for do equal to 61 mm (labeled as “O” in Table 3-2), whereas incipient necking at the hole tip was observed for do equal to 60 mm (labeled as “N” in Table 3-2), and a meridional crack was clearly developed for do equal to 59 mm (labeled as “F” in Table 3-2). Table 3-3 Experimental results and process parameters of the successful tests Edge radius, R (mm) d o (mm) d f D p (mm) h (mm) Bending ratio t o / R HER (d f / d o ) LFR (d f / d o,min ) 6 61 95.8 18.2 0.27 1.57 1.57 6 65 16.4 1.47 6 72 13.2 1.33 6 83 7.6 1.15 8 61 95.8 18.2 0.20 1.57 1.57 8 65 16.4 1.47 8 72 13.0 1.33 8 83 7.7 1.15 10 61 95.8 18.2 0.16 1.57 1.57 10 65 16.5 1.47 10 72 13.0 1.33 10 83 7.7 1.15 47.9 61 95.8 17.6 0.03 1.57 1.57 47.9 65 15.9 1.47 47.9 72 12.9 1.33 47.9 83 7.5 1.15 54 Hole flanging by conventional press working The mode of failure observed in all situations was the initiation of a neck at the edge of the flange, which developed until the material fractured. The onset of necking appeared preferentially oriented along the rolling direction, as shown in Figure 3.6. Table 3-3 presents the geometric parameters that were measured in the successful parts of the different pre cut holes. The final hole diameter df was assumed to be the punch diameter Dp, neglecting the slight springback experienced by the flange after removing the punch, as Dp was always the largest diameter that ran through the hole. The flange height h was measured at four different positions around the flange (approximately every 90°). The bending ratio, which is defined as the ratio of the initial sheet thickness to the punch edge radius (to/R), is also presented. This ratio quantifies the severity of local bending induced on the sheet thickness during the forming operation. With the exceptionof the hemispherical punch, the values were nearly identical for a given pre cut hole do when the bending ratios ranged from 0.16 to 0.27. 3.4. Formability analysis based on the LFR As mentioned in the introduction, the forming severity during conventional circular hole flanging is measured by the hole expansion ratio (HER), and the formability is quantified by the limiting forming ratio (LFR). These parameters are defined as follows: *+! = H    ∴ ()! ≔ H    , /2 (3-1) Figure 3.7 graphically depicts the information in Table 3-3, where flange height is dimensionalized by the final hole diameter df or the punch diameter Dp on the vertical axis versus the HER one the horizontal axis. The higher is the HER value, the higher is the non dimensional flange height. Note that all punch radii evolutions overlap, and the LFR or maximum HER, which is indicated by a dotted line, remained constant with R and was equal to 1.57, which indicates that the punch profile did not significantly influence successful flange formation, i.e., the flange formability. Similar conclusions have been noted by other authors, e.g., Leu et al., 1999 and Huang et al., 2001a. In their studies, they noted that the formability during hole flanging processes appeared to be unaffected by the punch profile. The reason for this insensibility will be 55 Analysis of single stage SPIF process applied to the hole flanging operation subsequently discussed. The dotted line joining the origin with the test indicates a suggested evolution as no data were available. For completeness, the dashed line in this process represents the physical limit for which the flange never becomes thicker on average, that is, the undeformed flange height (df - do)/2 should always be smaller than its average deformed height h. For a given HER, the flange height decreases with the bending ratio (to/R). As subsequently explained, in practice, due to the special geometry of the test, the tensile stress state of the edge of the pre cut hole produces a contraction of the flange, which is enhanced by the bending ratio and friction conditions of the punch. Smaller punch edge radii can reduce the contraction of the flange. Figure 3.7 Non-dimensional flange height (h/d f ) of successful conventional holeflanged sheets for different pre cut diameters 56 Hole flanging by conventional press working 3.5. Formability analysis based on the FLC Conventional FLC is currently the most useful tool during the designing stage of press working operations for evaluating the formability of a metal sheet. In general terms, safe operation is warranted when the strains in the sheet are well below the FLC. Figure 3.8 In-plane strains along the meridional direction for all successful hole flanged parts fabricated using different punches Figure 3.8 shows the in-plane strains along the meridional direction of finished flanges from all successful tests, i.e., the 61 mm, 65 mm, 72 mm and 83 mm pre cut holes using the cylindrical punches (6 mm, 8 mm, and 10 mm edge radius) and the hemispherical punch, respectively. The inspection directions were oriented along the rolling direction according to the preferential failure orientation. The principal strains, which were computed using ARGUS®, are represented by a solid line and cover points from the base to nearly the edge of the flange. The strain at the exact flange edge, which was obtained using direct measurement as indicated in section 0, is depicted with a star. The gap between the last point 63 Analysis of single stage SPIF process applied to the hole flanging operation 3.7. Conclusions The formability of the hole-flanged parts fabricated using press working was analysed via the LFR and the conventional FLC. The influence of the punch profile on the deformation process and flange failure was discussed. A series of conventional hole flanging tests using a 1.6 mm 7075-O aluminium alloy sheet were performed using three cylindrical punches with different edge radii and a hemispherical punch. The main conclusions of this study can be summarized as follows:  The failure mode prevailing at the formability limit was the initiation of necking at the edge of the flange.  For the successful parts fabricated with cylindrical punches with bending ratios that range from 0.16 to 0.27, the heights of the flanges were nearly identical for a given pre cut hole do. The hemispherical punch produced slightly lower flanges.  The LFR is a suitable measure of the formability in conventional hole flanging operations as this parameter directly quantifies the maximum stretching of the material at the hole tip, where failure occurs in practice. The LFR remained constant with R and was equal to 1.57.  In general, conventional FLC was not a suitable tool for analysing the formability of the material along the flange, with the exception of the flange tip.  The evolution of the in-plane principal strains on the outer surface of the flange exhibited a two-shaped curve. With the exception of the hemispherical punch, the strain curves for the pre cut holes near but below the formability limit were clearly above the FLC; however, failure did not occur. This event was observed in the area of the wall near the edge of the flange.  The bending effect seemed to be the main factor that controlled the apparent enhancement in the formability in this area, which was clearly influenced by the punch profile.  The punch edge radius primarily affected the final value of the meridional strain, i.e., the minor principal strains along the flange. A smaller punch edge radius induced a shift in the strain curves to the right in the principal 64 Hole flanging by conventional press working strain diagram. This shift was responsible for the lower value in the flange height exhibited by the semi-spherical punch.  The condition for the onset of necking at the flange tip was fulfilled at the intersection of the FLC with the line of pure uniaxial tension. The LFR value was associated with this intersection point, which explains why the LFR is a material constant that is independent of the punch profile and can be obtained as ()!=efg ( , ∗ ), where  , ∗ is the major strain at the onset of necking in uniaxial tension. In the next chapter, the formability and geometric capabilities of the hole flanging process by single point incremental forming (SPIF) in a single stage is discussed to critically evaluate the technological differences with respect to the conventional method by press working. 65 Analysis of single stage SPIF process applied to the hole flanging operation 4. H OLE FLANGING BY SINGLE STAGE SPIF he aim of the present chapter is to contribute to the understanding of the formability of metal sheets in the demanding operation of single-stage hole flanging by SPIF. Three important points need to be discussed: (1) the physical mechanisms that control the sheet failure during this process, (2) the maximum flange that can be successfully formed by SPIF in a single stage and (3), the geometrical capabilities of the process in terms of thickness distribution and flange shape. A series of experimental tests of AA7075-O metal sheets has been performed and analysed using three different tool radii (10, 8 and 6 mm) and two spindle speed values (0 and 1000 rpm). The surface roughness and thickness profile along the flanges on the final parts are also analysed. The forming forces are analysed to anticipate the failure of the metal sheet. 4.1. Experimental procedure The single-stage hole flanging tests were performed on a three-axis milling CNC machine EMCO VMC-200. The experimental setup is shown in Figure 4.1. The SPIF setup comprised a blank holder and a backing plate with a 100 mm diameter hole; both were fixed to the machine table via a rigid rig. The sheet metal blanks, which were initially without pre cut holes but electrochemically etched, were fixed upside down by the holder over the backing plate. The pre cut holes were milled and the edges were sanded to remove any burr. This sequence guaranties that the pre cut holes and tool are perfectly centre. Once the pre cut hole was finished and the tool was changed by the forming tool, the specimen was ready to be incrementally deformed, as shown in Figure 4.1 (a). To analyse the effect of the sheet thickness-to-tool radius ratio (to/R) on formability, three different radii were employed (R = 6, 8 and 10 mm). The forming tools are shown in Figure 4.2. T Holeflanging by single stage SPIF Figure 4. 1 flanging process by SPIF. (c) Scheme of the z Figure 4.1 (b) shows the scheme of the process. The test attempts to achieve the final theoretical geometry of the hole flanged sheet in a singl tool. In this case, a final cylindrical hole with an inner diameter programmed. (a) flanging by single stage SPIF 1 (a) Experimental setup for SPIF. (b) Scheme of the single flanging process by SPIF. (c) Scheme of the z frustum and cylindrical helical path Figure (b) shows the scheme of the process. The test attempts to achieve the final theoretical geometry of the hole flanged sheet in a singl tool. In this case, a final cylindrical hole with an inner diameter programmed. flanging by single stage SPIF (a) Experimental setup for SPIF. (b) Scheme of the single flanging process by SPIF. (c) Scheme of the z frustum and cylindrical helical path Figure 4.2 Formin g tools used for SPIF (b) shows the scheme of the process. The test attempts to achieve the final theoretical geometry of the hole flanged sheet in a singl tool. In this case, a final cylindrical hole with an inner diameter (b) (c) (a) Experimental setup for SPIF. (b) Scheme of the single flanging process by SPIF. (c) Scheme of the z - level trajectory for the conical frustum and cylindrical helical path g tools used for SPIF (b) shows the scheme of the process. The test attempts to achieve the final theoretical geometry of the hole flanged sheet in a singl tool. In this case, a final cylindrical hole with an inner diameter (a) Experimental setup for SPIF. (b) Scheme of the single - stage hole level trajectory for the conical frustum and cylindrical helical path g tools used for SPIF (b) shows the scheme of the process. The test attempts to achieve the final theoretical geometry of the hole flanged sheet in a singl e pass of the forming tool. In this case, a final cylindrical hole with an inner diameter df = 95.8 mm is 66 stage hole level trajectory for the conical (b) shows the scheme of the process. The test attempts to achieve the e pass of the forming = 95.8 mm is 67 Analysis of single stage SPIF process applied to the hole flanging operation A series of pre cut holes do was machined on the sheet blanks to obtain the formability limit of the single-stage hole flanging process by SPIF. This limit is defined as the maximum HER = df / do that ensures a complete flange that is free of defects, i.e., the LFR. The final flange height h and thickness are also measured to evaluate the geometric capabilities of the process. The forming trajectories to perform the hole flanging tests were modelled and simulated in CATIA V5® using a machining workbench. A virtual model of the experimental setup was employed to optimize the trajectories according to the desired final geometry of the flange. The tool trajectory was divided into two steps, as represented in Figure 4.1 (c). The tool follows a z-level anticlockwise path of concentric circles with a decreasing diameter to smoothly adjust the sheet to the backing plate. The second step is a cylindrical helical anti-clockwise path. A step down of 0.2 mm/turn is utilized in both steps. Note that a simpler alternative to the second step is to use a trajectory of successive circular paths of constant diameters at different depths. However, a drawback of this alternative is that the vertical force showed a significant peak when the tool went down between two circular paths, which increased the probability of premature failure in the sheet at this location. Two tool rotation conditions were tested:0 rpm and 1000 rpm clockwise. The condition of 0 rpm represents a tool that cannot rotate at all, i.e., a locked tool. The feed rate was set to 1000 mm/min. To minimize the friction effects, a special lubricant for metal forming applications, Castrol Iloform TDN81, was continuously applied between the forming tool and the sheet during deformation. A Kistler® dynamometer was placed between the machine-table and the SPIF frame (see Figure 4.1(a)) to measure the forming forces in the X-, Y- and Z-axes during the tests. The forces were acquired using the commercial software LabVIEW®. As a reference, the typical duration of a SPIF test was approximately 40-45 minutes for the successful specimens and approximately 15-20 minutes for the fractured specimens. Approximately 15 minutes were initially taken for milling and deburring the hole. To study the deformation and failure mechanisms, the strains at the outer sheet surface were obtained by measuring with the optical 3D forming analysis system ARGUS®. To analyse the geometric capability of the process, finished flanges of successful specimens were cut to measure the final height, the sheet thickness distribution along the produced flange and the surface roughness. 68 Hole-flanging by single stage SPIF 4.2. Results According to the experimental procedure, six series of tests were performed for three different tool radii (R = 6, 8 and 10 mm) and two spindle speeds (0 and 1000 rpm clockwise). Table 4-1 lists the values of the pre cut hole diameters do and their results. The label “O” represents a successful test in terms of obtaining a complete hole-flanged part with no defects, “X” indicates fractured specimens, and “D” denotes specimens that have a complete flange but exhibit surface defects. To determine the forming capability for each test condition, a series of increasing values of do was tested to achieve a successful flange. Intermediate pre cut holes were tested until a difference of less than or equal to 1 mm in diameter between successful tests and failed tests was reached. Note that the failed test (X and/or D) closest to the first successful test was repeated at least twice in all cases to confirm the failure. In the case of the successful tests, only a few tests in the frontier of formability were repeated. In both cases, the results showed a high repeatability. Table 4-1 Series of SPIF tests Tool R (mm) Spindle (rpm) Initial pre cut diameter of the hole, do (mm) 55 56 57 57.5 58 59 60 61 62.5 63.5 64.5 65 70 82 6 0 X X X X O O O O 6 1000 X O O 8 0 X X O O O 8 1000 X O O 10 0 X X D O O O O O 10 1000 X X O O Figure 4.3 shows pictures of the closest successful-failed test pair for every tool radius at 0 rpm. The failed tests were interrupted as soon as the initiation of a crack was observed. Three deformation zones can be distinguished on the inprocess workpiece: the cylindrical flange that is already formed, the curved sheet in contact with the tool in the process of forming, and the still-flat sheet that is Analysis of single stage SPIF process applied to the hole flanging operation being radially stretched. The latter is essential the larger is this zone, the greater is the deformation resistance. Similar findings are also observed at 1000 rpm. Figure 4.3 Pictures of the closest success Analysis of single stage SPIF process applied to the hole flanging operation being radially stretched. The latter is essential the larger is this zone, the greater is the deformation resistance. Similar findings are also observed at 1000 rpm. Pictures of the closest success Analysis of single stage SPIF process applied to the hole flanging operation being radially stretched. The latter is essential the larger is this zone, the greater is the deformation resistance. Similar findings are also observed at 1000 rpm. Pictures of the closest success - failed te Details of the deformed sheets Analysis of single stage SPIF process applied to the hole flanging operation being radially stretched. The latter is essential to the success of the operation as the larger is this zone, the greater is the deformation resistance. Similar findings failed te st pair with a spindle speed of 0 rpm. Details of the deformed sheets Analysis of single stage SPIF process applied to the hole flanging operation to the success of the operation as the larger is this zone, the greater is the deformation resistance. Similar findings st pair with a spindle speed of 0 rpm. Details of the deformed sheets 69 Analysis of single stage SPIF process applied to the hole flanging operation to the success of the operation as the larger is this zone, the greater is the deformation resistance. Similar findings st pair with a spindle speed of 0 rpm. 70 Hole-flanging by single stage SPIF Table 4-2 Experimental results and process parameters in successful tests with a tool rotation of 1000 rpm Tool R d o d f h HER LFR to/R TR (mm) (mm) (mm) (mm) (d f / d ) (HER max ) (    3 ⁄ ) 6 64.5 94.3 22.3 1.46 1.46 0.27 0.56 6 65 94.3 21.4 1.45 0.27 0.58 8 60 94.8 26.4 1.58 1.58 0.20 0.54 8 61 94.7 24.6 1.55 0.20 0.56 10 57 94.8 26.6 1.66 1.66 0.16 0.57 10 57.5 94.8 25.5 1.65 0.16 0.59 Table 4-3 Experimental results and process parameters in successful tests with a tool rotation of 0 rpm Tool R (mm) d o (mm) d f (mm) h (mm) HER (d f / d o ) LFR (HER max ) t o /R TR ( i  / i k ) 6 64.5 94.4 23.6 1.46 1.46 0.27 0.53 6 65 94.4 22 1.45 0.27 0.56 6 70 94.4 16 1.35 0.27 0.66 6 82 95.2 8.8 1.16 0.27 0.70 8 61 94.8 24.8 1.55 1.55 0.20 0.56 8 62.5 94.7 22 1.52 0.20 0.61 8 65 94.7 19 1.46 0.20 0.66 10 57.5 94.8 26.8 1.65 1.65 0.16 0.56 10 58 94.6 25.5 1.63 0.16 0.58 10 61 94.8 21 1.55 0.16 0.66 10 65 94.8 17.8 1.46 0.16 0.71 10 82 95.3 8.3 1.16 0.16 0.75 71 Analysis of single stage SPIF process applied to the hole flanging operation Table 4-2 and Table 4-3 present different geometric and process parameters measured in the successful parts for different pre cut holes for spindle speeds of 1000 rpm and 0 rpm, respectively. The flange height h and the final hole diameter df were measured at four different positions around the flange (approximately every 90º). A Mitutoyo digital calliper with a resolution of 0.01 mm was employed. The tables list the mean value of h and df; their standard deviation was slower than 0.15 mm in both cases. Note that df slightly differs from its theoretical value of 95.8 mm. Apart from the intrinsic experimental scatter, these differences seem to be primarily attributed to the elastic deflection of the tool and the elastic springback of the sheet observed in the tests. As a result, a slight trend to obtain smaller final holes with a smaller tool radius, i.e., with more flexible tools, is observed. The tables also present four different process parameters that are used to analyse the process. The HER and the LFR were defined in the introduction. The bending ratio is defined as the ratio of the initial sheet thickness-to-tool radius (to/R), which quantifies the severity of local bending induced by the forming tool. The average thickness ratio is defined as the ratio of the average thickness   of the final flange to the initial sheet thickness (  /  ). The interest in these parameters to explain the deformation process is discussed in the following sections. A comparison of the results at 1000 and 0 rpm reveal that the higher spindle speed slightly decreases the minimum pre cut hole allowed and seems to yield marginally shorter flanges for the same pre cut hole. This fact should be considered with caution since these differences are small in both cases and may be influenced by the intrinsic dispersion in the experimental tests and measures. As will be subsequently discussed, increasing the tool rotation provides clear beneficial effects, such as a more uniform thickness distribution in the final flange, a reduction in the roughness at the surface in contact with the tool and a slight decrease in the forming force. 4.3. Strain analysis and failure mechanism The strain analysis of the outer sheet surface enables an understanding of the physical mechanism that controls the sheet failure. By comparing the strain distribution evolution from successful tests to failed tests, this mechanism can be clearly observed. Figure 4.4 shows the strain evolution in the flange for specimens with different pre cut holes (58, 57.5, 57 and 56 mm diameters), formed with a 10 mm tool radius at 72 Hole-flanging by single stage SPIF 0 rpm. The principal strains were measured along the flange in line a-b-c-d (see the left legend in Figure 4.3). The strain evolutions in the FLD describe a loop, beginning from the bottom of the flange (point a), reaching a maximum value near the midway (point b) and ending at the flange edge (point d). Strains measured with ARGUS® are identified with closed symbols, whereas strains calculated at the flange edge (point d) are represented by open stars. In the failed specimens, the fracture strains measured on the crack lips are represented by open symbols. In all cases, the strains are above the FLC for necking obtained by the Nakazima tests, which indicates that necking is clearly postponed or suppressed. This inhibition of necking has also been reported by Silva et al.,2011 and Centeno et al.,2014 in SPIF processes. The smaller is the pre cut hole diameter, the higher is the major strain in the central region of the flange (point b). Figure 4.4 FLD with the evolution of strain paths along the final flange for different pre cut hole diameters for a R10 tool radius at 0 rpm In this series, the holed specimens with diameters of 58 and 57.5 mm were successfully formed (labelled with “O” in Table 4-1), whereas the holed specimen with a diameter of 56 mm fractured before completing the flange (labelled with “X”). As expected, the former exhibited strain loops below the FFL, whereas the latter clearly crossed the FFL. Special attention should be paid to the specimen 79 Analysis of single stage SPIF process applied to the hole flanging operation Figure 4.10 Non-dimensional flange height (h/d f ) of successful hole-flanged sheets for different pre cut diameters at 0 rpm A direct parameter that quantifies the stretching that the flange underwent is the flange heighth presented in Table 4-3. Figure 4.10 depicts the evolution of the non-dimensional flange height h/df versus the HER. For a given HER (for instance, the vertical dotted line), the non-dimensional flange height h/df increases as the tool radius decreases, which shows agreement with the expected trend and indicates that the formability of the flange increases as the bending ratio (to/R) increases. 80 Hole-flanging by single stage SPIF Note that the non-dimensional flange height h/df in Figure 4.10 has been extrapolated to zero for a pre cut hole d0 that is equal to the final hole df (HER = 1). For large pre cut holes (typically, HER<1.2), the flange height is practically independent of the tool radius and no gain in formability is obtained from reducing the tool radius. As indicated in the previous chapter, the physical limit for the flange height is depicted (dashed line). According to these results, a more appropriate parameter for measuring the formability of the flange is the thickness ratio, which is obtained by assuming volume constancy during the flange deformation. This approach yields the formula π  ℎ= π 4(   −   )  (4-1) where  m is the average thickness in the finished flange. Note that this equation realistically combines the axial, the circumferential and the thickness deformation undergone by the flange. Thus, the average thickness ratio (TR) may be written as follows: n! =     =    −    4   ℎ → n! = 1 − 1 *+!  ⁄ 4 ( ℎ   ⁄ ) (4-2) As shown in Eq. (4-2), for a given HER, the higher is the flange, the lower is the average thickness ratio. The experimental values of TR are shown in Table 4-3. For a given HER, TR decreases as the bending ratio (to/R) increases. As expected, in practice, smaller tools are able to further stretch the material of the flange provided that the initial pre cut hole is above the critical hole established by the minimum radius tool. A similar trend is also observed in tests at 1000 rpm, as shown in Table 4-2. An exception is obtained for the tool with an 8 mm radius, which yield a TR value that is notably smaller than that obtained with the 6 mm radius tool. Analysis of single stage SPIF process applied to the hole flanging operation 4.5. shape The thickness profil (magnification, x20) and commercial image processing software. presents the thickness profile along the flange in the rolling direction obtained in two hole flanging tests at 0 and 1000 rpm. No significant differences were observed in the shape of the pr correspond to tests performed with a 6 mm radius tool in blanks with a 64.5 mm diameter hole. Figure 4.11 Sheet thickness along the flange formed with spind The thickness reduction presents a wavy profile along the flange height, with a maximum reduction of 62% (0 rpm) and 53% (1000 rpm) around the middle of the flange (point hole flanging noted by Cui distribution by forming a successive conical frustum of an increasing angle to 90º. Although the thickness distribution provided by SPIF can be improved using an appropriate multi al., 2014, assuming its geometrical restrictions, a single functional flanges in considerably less time. As deduced from Analysis of single stage SPIF process applied to the hole flanging operation 4.5. Thickness shape The thickness profil es were measured with a Nikon SMZ800 (magnification, x20) and commercial image processing software. presents the thickness profile along the flange in the rolling direction obtained in two hole flanging tests at 0 and 1000 rpm. No significant differences were observed in the shape of the pr correspond to tests performed with a 6 mm radius tool in blanks with a 64.5 mm diameter hole. Sheet thickness along the flange formed with spind The thickness reduction presents a wavy profile along the flange height, with a maximum reduction of 62% (0 rpm) and 53% (1000 rpm) around the middle of the flange (point b in Figure hole flanging noted by Cui distribution by forming a successive conical frustum of an increasing angle to 90º. Although the thickness distribution provided by SPIF can be improved using an appropriate multi - stage strategy, as reporte 2014, assuming its geometrical restrictions, a single functional flanges in considerably less time. from Figure 4 . Analysis of single stage SPIF process applied to the hole flanging operation Thickness , surface roughness es were measured with a Nikon SMZ800 (magnification, x20) and commercial image processing software. presents the thickness profile along the flange in the rolling direction obtained in two hole flanging tests at 0 and 1000 rpm. No significant differences were observed in the shape of the pr ofiles measured in other directions. The results correspond to tests performed with a 6 mm radius tool in blanks with a 64.5 mm Sheet thickness along the flange formed with spind rpm The thickness reduction presents a wavy profile along the flange height, with a maximum reduction of 62% (0 rpm) and 53% (1000 rpm) around the middle of Figure 4.3 (upper)). The experimental results of multi hole flanging noted by Cui et al., 2010 present a similar evolution in the thickness distribution by forming a successive conical frustum of an increasing angle to 90º. Although the thickness distribution provided by SPIF can be improved using an stage strategy, as reporte 2014, assuming its geometrical restrictions, a single functional flanges in considerably less time. . 11 , a higher spindle speed yields a final flange with a Analysis of single stage SPIF process applied to the hole flanging operation surface roughness es were measured with a Nikon SMZ800 (magnification, x20) and commercial image processing software. presents the thickness profile along the flange in the rolling direction obtained in two hole flanging tests at 0 and 1000 rpm. No significant differences were ofiles measured in other directions. The results correspond to tests performed with a 6 mm radius tool in blanks with a 64.5 mm Sheet thickness along the flange formed with spind rpm The thickness reduction presents a wavy profile along the flange height, with a maximum reduction of 62% (0 rpm) and 53% (1000 rpm) around the middle of (upper)). The experimental results of multi 2010 present a similar evolution in the thickness distribution by forming a successive conical frustum of an increasing angle to 90º. Although the thickness distribution provided by SPIF can be improved using an stage strategy, as reporte d by Skjoedt 2014, assuming its geometrical restrictions, a single - stage strategy may provide functional flanges in considerably less time. , a higher spindle speed yields a final flange with a Analysis of single stage SPIF process applied to the hole flanging operation surface roughness and flange es were measured with a Nikon SMZ800 (magnification, x20) and commercial image processing software. Figure presents the thickness profile along the flange in the rolling direction obtained in two hole flanging tests at 0 and 1000 rpm. No significant differences were ofiles measured in other directions. The results correspond to tests performed with a 6 mm radius tool in blanks with a 64.5 mm Sheet thickness along the flange formed with spind le speeds of 0 and 1000 The thickness reduction presents a wavy profile along the flange height, with a maximum reduction of 62% (0 rpm) and 53% (1000 rpm) around the middle of (upper)). The experimental results of multi 2010 present a similar evolution in the thickness distribution by forming a successive conical frustum of an increasing angle to 90º. Although the thickness distribution provided by SPIF can be improved using an d by Skjoedt et al., 2010 and Mirnia stage strategy may provide , a higher spindle speed yields a final flange with a 81 Analysis of single stage SPIF process applied to the hole flanging operation and flange es were measured with a Nikon SMZ800 microscope Figure 4.11 presents the thickness profile along the flange in the rolling direction obtained in two hole flanging tests at 0 and 1000 rpm. No significant differences were ofiles measured in other directions. The results correspond to tests performed with a 6 mm radius tool in blanks with a 64.5 mm le speeds of 0 and 1000 The thickness reduction presents a wavy profile along the flange height, with a maximum reduction of 62% (0 rpm) and 53% (1000 rpm) around the middle of (upper)). The experimental results of multi -stage 2010 present a similar evolution in the thickness distribution by forming a successive conical frustum of an increasing angle to 90º. Although the thickness distribution provided by SPIF can be improved using an 2010 and Mirnia et stage strategy may provide , a higher spindle speed yields a final flange with a Holeflanging by single stage SPIF slightly more uniform thickness profile, which reduces the risk of sheet failure. Durante et al., temperature . The lower friction reduces the forming force and stretching stresses, which promotes a lower thickness reduction in the area that is being deformed. Figure 4. 12 flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool Another beneficial effect of increasing the tool rotation is the lower Flange deformed a Flange deformed at 1000 rpm, rolling direction ( flanging by single stage SPIF slightly more uniform thickness profile, which reduces the risk of sheet failure. et al., 2009 attributed this effect to friction reduction and increased . The lower friction reduces the forming force and stretching stresses, which promotes a lower thickness reduction in the area that is being deformed. 12 Roughness profile Ra of a blank sheet (0° and 90° RD) and deformed flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool radius and 65 mm initial hole diameter Another beneficial effect of increasing the tool rotation is the lower Blank, transverse direction ( Blank, rolling direction ( Flange deformed a Flange deformed at 1000 rpm, rolling direction ( flanging by single stage SPIF slightly more uniform thickness profile, which reduces the risk of sheet failure. 2009 attributed this effect to friction reduction and increased . The lower friction reduces the forming force and stretching stresses, which promotes a lower thickness reduction in the area that is being deformed. Roughness profile Ra of a blank sheet (0° and 90° RD) and deformed flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool radius and 65 mm initial hole diameter Another beneficial effect of increasing the tool rotation is the lower Blank, transverse direction ( Blank, rolling direction ( Flange deformed a 0 rpm, rolling direction ( Flange deformed at 1000 rpm, rolling direction ( slightly more uniform thickness profile, which reduces the risk of sheet failure. 2009 attributed this effect to friction reduction and increased . The lower friction reduces the forming force and stretching stresses, which promotes a lower thickness reduction in the area that is being deformed. Roughness profile Ra of a blank sheet (0° and 90° RD) and deformed flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool radius and 65 mm initial hole diameter Another beneficial effect of increasing the tool rotation is the lower Blank, transverse direction ( Ra =0.29 μ Blank, rolling direction ( Ra=0.07μ m) rpm, rolling direction ( Ra Flange deformed at 1000 rpm, rolling direction ( Ra slightly more uniform thickness profile, which reduces the risk of sheet failure. 2009 attributed this effect to friction reduction and increased . The lower friction reduces the forming force and stretching stresses, which promotes a lower thickness reduction in the area that is being deformed. Roughness profile Ra of a blank sheet (0° and 90° RD) and deformed flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool radius and 65 mm initial hole diameter Another beneficial effect of increasing the tool rotation is the lower =0.29 μ m) m) Ra =0.34 μm) Ra =0.17 μm) 82 slightly more uniform thickness profile, which reduces the risk of sheet failure. 2009 attributed this effect to friction reduction and increased . The lower friction reduces the forming force and stretching stresses, which promotes a lower thickness reduction in the area that is being deformed. Roughness profile Ra of a blank sheet (0° and 90° RD) and deformed flange (0 and 1000 rpm),which correspond to a specimen tested with a 6mm tool Another beneficial effect of increasing the tool rotation is the lower roughness 83 Analysis of single stage SPIF process applied to the hole flanging operation observed in the inner sheet surface in contact with the tool. Figure 4.12 shows the roughness profile Ra that is measured in the blank prior to forming, in the rolling and transverse directions and in the final flange at 0 and 1000 rpm. A Mahr® M1 perthometer was employed. Ra is reduced by almost 50% when the speed rotation is increased from 0 rpm (0.34 µm) to 1000 rpm (0.17 µm). Note that the absolute values of Ra are highly dependent on the step down, which is fixed to 0.2 mm in all tests in this study. This improvement of superficial quality has also been attributed to friction reduction by Durante et al.,2009. Numerous authors have recently investigated new tools and techniques to reduce friction in SPIF to enhance the formability and surface quality, e.g., using laser surface texturing tools, as proposed by Xu et al., 2013, or oblique roller ball tools, as suggested by Lu et al.,2014. The final shape must be considered. In Figure 4.13, two sections are superposed, red and black. The section in red is the section of the theoretical shell that is used to program the theoretical flange profile by CATIA® V5 to a final hole flange with a diameter of 95.8 mm. The black section is the result of a hole flanging by SPIF using a tool diameter of 16 mm and a pre cut hole with a diameter of 61 mm. The scanned image and section were performed with the 3D Geomagic Capture® scanner with a precision between 0.060 and 0.118 mm depending on the capture distance. Due to the non symmetry of the forming process (the tool is shaping on one side while nothing occurs on the opposite side) and the stiffness of the tool, which acts as a small cantilever beam (stiffness is especially low for small tool diameters), achieving a theoretical final diameter is challenging. Figure 4.13 Section of specimen R8 with a pre cut hole with a61mm diameter, scanned and compared with the theoretical profile. Although low springback effects were also observed, the tendency seemed to be regular and only a wider tool trajectory design would be sufficient for achieving a 95.8 mm diameter hole flange. 84 Hole-flanging by single stage SPIF 4.6. Forming Forces According to part geometry and tool trajectories, the two main force components involved in the process are the vertical force Fz and the horizontal force Fh. The former, which is directly measured by the dynamometer, is primarily related to the stiffness in the axial direction of the flange and the friction force in the tool due to the material flow from the flat part towards the flange in the wall. The latter is related to the circumferential deformation of the flat part and the friction force in the direction of the tool feed, which can be calculated from the measured forces Fx and Fy as ) Q = p ) 1  + ) q  (4-3) Figure 4.14 Horizontal and vertical force components vs. tool vertical displacement for successful tests at different spindle speeds. 85 Analysis of single stage SPIF process applied to the hole flanging operation Figure 4.14 compares the typical evolution of Fh and Fz in two successful tests with a tool radius of 6 mm at 0 and 1000 rpm. As expected, smaller forming forces are required at 1000 rpm. In both cases, the vertical force Fz predominates over Fh at early stages of the forming process. This force rapidly increases to a maximum value and decreases to zero at the end of the test. The horizontal force shows a similar evolution but subsequently reaches the maximum value. A nonzero value is obtained at the end of the test due to the springback force exerted by the formed flange over the tool. In both cases, the transition between the z-level and the helical path of the tool is identified. A comparison of the evolution of the horizontal force-to-vertical force ratio (Fh/Fz) during the test for a successful specimen and a failed specimen is interesting. Figure 4.15 shows this ratio for the successful tests discussed in Figure 4.14 (65 mm diameter pre cut hole) and their nearest failed tests (63.5 mm diameter pre cut hole). Excluding the initial z-level phase, in successful tests, the Fh/Fz ratio monotonically increases during the helical phase until the flange is finished. However, in failed tests, the Fh/Fz ratio shows a maximum value and then decreases until failure occurs. This evolution was observed in all tests analysed in this study. Figure 4.15 Horizontal-to-vertical forces ratio (F h / F z ) for successful and failed tests at different spindle speeds 86 Hole-flanging by single stage SPIF The explanation for this increasing F h /F z ratio in successful tests is based on the fact that the material in the wall is more resistant than the still non deformed material in the flat part (below the forming tool), yielding an appropriate material flow to form the flange. However, the appearance of a maximum and the subsequent decrease in the F h /F z ratio indicates that the deformation in the wall is more intense than in the flat part, giving rise to a failure by excessive stretching of the wall. Apart from the physical explanation, note that the appearance of this maximum in the forces occurs in advance of the final failure. Therefore, monitoring this force ratio during the process may help to anticipate the sheet failure and control the process in practice. To study the tool punch radius effect in forming forces, Figure 4.16 compares the three axial forming forces of the tool radii 6 mm, 8 mm and 10 mm for the same initial 65 mm diameter pre cut hole and 0 rpm spindle speed. Figure 4.16 Vertical Forces for an initial pre cut hole with a 65 mm diameter and punch tool radii of 6, 8 and 10 mm Although the total amount of material to be conformed was the same for all three tests, the larger radius of the punch allowed the material to flow more easily from 87 Analysis of single stage SPIF process applied to the hole flanging operation the flat part to the wall than the smaller radius; thus, the axial forces were lower. Reducing the friction coefficient by increasing the spindle speed can balance the punch radius effect. Figure 4.17 shows the radial force evolution for a punch tool radius of 10 mm and 8 mm at 0 rpm and 6 mm with and without a spindle speed for the same initial pre cut hole of 65 mm diameter. Contrary to the axial force evolution, a larger tool radius produces higher radial forces due to the amount of material involved in the forming process. Figure 4.17 Radial forces for an initial pre cut hole with a 65 mm diameter and punch tool radii of 6, 8 and 10 mm 4.7. Conclusions The capability of SPIF to successfully perform hole flanging operations using a single-stage strategy has been analysed. The material is a 7075-O aluminium alloy sheet with a thickness of 1.6 mm. A series of sheet metal blanks with different pre cut holes has been incrementally deformed using three different hemispherical forming tools with radii of 10 mm, 8 mm and 6 mm and two spindle speeds, 0 and 1000 rpm clockwise. The main conclusions can be summarized as follows: 88 Hole-flanging by single stage SPIF  The dominant failure mechanism in this process has been determined. As discussed, the sheet failure is produced by biaxial stretching near the midway of the deformed flange instead of uniaxial tension at the hole edge, as usually observed in conventional hole flanging operations.  Typical for SPIF, the bending ratio to/R has an important role in the failure mode. For the smaller tools (6 mm and 8 mm radii), the mechanism of failure is ductile fracture with a complete suppression of local necking. For the larger tool (10 mm radius), the observed mechanism is ductile fracture with postponed necking at the very early stage.  A higher limiting forming ratio LFR obtained is 1.66 for the larger tool radius. For the tools in this study, the LFR increases as the tool radius increases.  As failure in incremental hole flanging occurs via severe stretching and thinning of the flange, the non-dimensional flange height (h/df) and the average thickness ratio ( 3 ⁄) seem to be more appropriate parameters for evaluating the geometric capability of the process. Thus, as the tool radius decreases, h/df increases and  3 ⁄ decreases for a given initial pre cut hole. This finding shows agreement with the expected trend in SPIF, which indicates that the capability of forming a flange increases as the bending ratio (to/R) increases.  The use of spindle speed does not significantly increase the formability in terms of the minimum pre cut hole to be successfully machined. However, spindle speed provides other interesting advantages, such as better surface quality, a more regular thickness profile along the flange and a reduction in forming forces. Therefore, the use of spindle speed is highly recommended in this demanding single-stage process.  To increase the dimensional accuracy and due to low springback effects, initial theoretical tool trajectories need to be modified.  The evolution of forming forces during the process, particularly the ratio of the horizontal component to the vertical component, can be useful in predicting sheet failure in practice. In the next chapter, a comparison between the conventional hole flanging process and the incremental hole flanging process will be presented. 95 Analysis of single stage SPIF process applied to the hole flanging operation 5.3.2. Analysis in terms of the LFR As discussed in chapter 3, the Limiting Forming Ratio (LFR) (see Eq. (1-1)) is the parameter that is traditionally employed to quantify the formability in hole flanging operations. Table 5-2 summarizes the LFR values obtained in both singlestage SPIF and press working process using different tool radii (R6, R8, R10 and R47.9 (only press working)). Table 5-2 Summary of LFR values in hole flanging by press working and single-stage SPIF and their corresponding non-dimensional height and average thickness of the flange for these specimens Hole-flanging Tool R6 Tool R8 Single-stage SPIF Spindle (rpm) d 0 mm LFR ℎ   ⁄     ⁄ d 0 mm LFR ℎ   ⁄     ⁄ 0 64.5 1.46 0.25 0.53 61 1.55 0.26 0.56 1000 64.5 1.46 0.24 0.56 60 1.58 0.28 0.54 Press-working 61 1.57 0.19 0.78 61 1.57 0.19 0.78 Hole-flanging Tool R10 Tool R47.9 Single-stage SPIF Spindle (rpm) d 0 mm LFR ℎ   ⁄     ⁄ d 0 mm LFR ℎ   ⁄     ⁄ 0 57.5 1.65 0.28 0.56 - - - - 1000 57 1.66 0.28 0.57 - - - - Press-working 61 1.57 0.19 0.78 61 1,57 0,18 0,81 Three facts can be clearly highlighted. The first fact is the insensitivity of the LFR to the punch edge radius in conventional press working, as discussed in detail in chapter 3, where the LFR is 1.57. The second fact, which was analysed in chapter 4, is the relative increase in the LFR value as the tool radius increases in singlestage SPIF. The maximum LFR obtained in this case is 1.66. The third fact is that the gain in formability measured via the LFR is not significant in practical terms. Both processes present an approximate LFR value of 1.6. This fact highlights the ability of the LFR to compare the formability between both processes. As previously discussed in chapter 4, this parameter considers the maximum circumferential strain at the hole edge, which is consistent with the failure site in a conventional hole flanging operation but far from the failure location observed in SPIF. Two parameters were proposed to measure the 96 Conventional vs. Incremental hole-flanging formability of the flange: the non-dimensional height of the flange (ℎ   ⁄), as a measure of the stretching undergone by the material in the flange in the meridional direction, and the average thickness ratio ( 3 ⁄), as a measure of the reduction in thickness undergone by the material in the flange (see Eq. (4-2)). Thus, the higher is the value of ℎ   ⁄ and the lower is the  3 ⁄, the higher is the formability. Figure 5.5 Evolution of the non-dimensional flange height (ℎ   ⁄) vs. HER of successful hole flanging tests manufactured by press working and single-stage SPIF at 0 rpm These two parameters are displayed in Table 5-2. As expected, both show insensitivity to the punch edge radius in conventional press working, with values 97 Analysis of single stage SPIF process applied to the hole flanging operation of approximately 0.19 and 0.78 for non-dimensional height and average thickness ratio, respectively. However, these values significantly different in SPIF, where values of ℎ   ⁄, which range from 0.25-0.28, and values of 3 ⁄, whichrange from 0.53-0.56, are typically observed at 0 rpm. Similar values are obtained at 1000 rpm, that is, ℎ   ⁄values from 0.24-0.28 and  3 ⁄values from 0.54-0.57. Figure 5.6 Evolution of the average thickness ratio ( 3 ⁄) vs. HER of successful hole flanging tests manufactured by press working and single-stage SPIF at 0 rpm Figure 5.5 and Figure 5.6 graphically depict the differences between press working and single-stage SPIF at 0 rpm. All successful hole flanging tests manufactured by both processes are compiled. The experimentally achieved Conventional vs. Incremental hole ranges of ℎ ⁄ evolution, the results concluded that single enhancement in formability quantified at approximately 40% in terms of relative flange heigh approximately 30% in terms of relative thickness reduction ( improvement does not manifest when using traditional flangeability, which reflects the fact that this parameter does not reproduce the mechanism of 5.4. The final shape of the flange is another important factor to compare the feasibility of both processes. The geometrical feature of flanges obtained in tests with a 61 mm pre cut hole diameter and a 8 m rpm are analysed. Figure 5.7 details both specimens before they were processed. They were cut along the meridional direc scanned using the scanner model 3D Geomagic Capture to shape and thickness evolution. Figure 5.7 Flanged specimens with a pre cut hole with a diameter of 61 mm and 8 mm Conventional vs. Incremental hole   ⁄ ,  3 ⁄ and evolution, the results concluded that single enhancement in formability quantified at approximately 40% in terms of relative flange heigh approximately 30% in terms of relative thickness reduction ( improvement does not manifest when using traditional flangeability, which reflects the fact that this parameter does not reproduce the mechanism of failure in hole flanging by SPIF. 5.4. Geometrical features The final shape of the flange is another important factor to compare the feasibility of both processes. The geometrical feature of flanges obtained in tests with a 61 mm pre cut hole diameter and a 8 m rpm are analysed. details both specimens before they were processed. They were cut along the meridional direc tion of the flange and polished. The flange profiles were using the scanner model 3D Geomagic Capture to shape and thickness evolution. Flanged specimens with a pre cut hole with a diameter of 61 mm and 8 mm tool radius by SPIF at 0 rpm (a) and press working (b) Conventional vs. Incremental hole -flanging and LFR versus evolution, the results concluded that single enhancement in formability regarding the press working, which quantified at approximately 40% in terms of relative flange heigh approximately 30% in terms of relative thickness reduction ( improvement does not manifest when using traditional flangeability, which reflects the fact that this parameter does not reproduce the failure in hole flanging by SPIF. Geometrical features The final shape of the flange is another important factor to compare the feasibility of both processes. The geometrical feature of flanges obtained in tests with a 61 mm pre cut hole diameter and a 8 m m tool radius for press working and SPIF at 0 details both specimens before they were processed. They were cut tion of the flange and polished. The flange profiles were using the scanner model 3D Geomagic Capture to shape and thickness evolution. Flanged specimens with a pre cut hole with a diameter of 61 mm and 8 mm tool radius by SPIF at 0 rpm (a) and press working (b) versus HER are visualized. Comparing their evolution, the results concluded that single - stage SPIF exhibits a significant regarding the press working, which quantified at approximately 40% in terms of relative flange heigh approximately 30% in terms of relative thickness reduction ( improvement does not manifest when using traditional flangeability, which reflects the fact that this parameter does not reproduce the failure in hole flanging by SPIF. Geometrical features The final shape of the flange is another important factor to compare the feasibility of both processes. The geometrical feature of flanges obtained in tests with a 61 m tool radius for press working and SPIF at 0 details both specimens before they were processed. They were cut tion of the flange and polished. The flange profiles were using the scanner model 3D Geomagic Capture to Flanged specimens with a pre cut hole with a diameter of 61 mm and 8 mm tool radius by SPIF at 0 rpm (a) and press working (b) are visualized. Comparing their stage SPIF exhibits a significant regarding the press working, which is approximately quantified at approximately 40% in terms of relative flange heigh t ( approximately 30% in terms of relative thickness reduction (   improvement does not manifest when using traditional LFR as a measure of flangeability, which reflects the fact that this parameter does not reproduce the The final shape of the flange is another important factor to compare the feasibility of both processes. The geometrical feature of flanges obtained in tests with a 61 m tool radius for press working and SPIF at 0 details both specimens before they were processed. They were cut tion of the flange and polished. The flange profiles were using the scanner model 3D Geomagic Capture to measure their local Flanged specimens with a pre cut hole with a diameter of 61 mm and 8 mm tool radius by SPIF at 0 rpm (a) and press working (b) 98 are visualized. Comparing their stage SPIF exhibits a significant is approximately t ( ℎ   ⁄) and    3 ⁄). This as a measure of flangeability, which reflects the fact that this parameter does not reproduce the The final shape of the flange is another important factor to compare the feasibility of both processes. The geometrical feature of flanges obtained in tests with a 61 m tool radius for press working and SPIF at 0 details both specimens before they were processed. They were cut tion of the flange and polished. The flange profiles were measure their local Flanged specimens with a pre cut hole with a diameter of 61 mm and 8 mm 99 Analysis of single stage SPIF process applied to the hole flanging operation Figure 5.8 displays the flange profile of both specimens. The continuous red line represents the theoretical internal profile of the flange. The distortion is visible in both cases. In the SPIF case, the flange accurately reproduce the theoretical shape at the base and shows a slight inward distortion at the wall. This wall distortion is primarily attributed to the springback of the flange but can also be influenced by the tool deflexion that is exhibited to a certain extent in this process. In the case of press working, although distortion of the wall is similar or even seems to improve regarding the incremental case, the flange shows an important springback at the base. This springback is directly related to the blank-holder force and clamping condition during the tests. This distortion may be improved by increasing the blank-holder force and modifying the clamping system. However, this analysis is beyond the scope of this study, which was originally focused on the comparative evaluation of the formability between both hole flanging processes. Figure 5.8 Comparison of flange profile from specimens with a 61mm pre cut hole diameter and 8 mm tool radius by SPIF at 0 rpm (left) and press working (right) The thickness distribution versus the flange height in both tests are shown in Figure 5.9. As discussed in section 0, the thickness distribution in the incremental flange is considerably more irregular than that in press working, as a result of the high strain gradient exhibited in the meridional direction in this process (see Figure 5.2). The large thickness reduction obtained around the middle of the flange is also observed. According to the Figure 5.9, the local and incremental deformation induced in the single-stage SPIF can intensively stretch the material, which enables higher and thinner flanges for a given HER parameter. Conversely, the conventional hole flanging process provides shorter but thicker (more robust) flanges for an identical specimen. Thus, the practical application and service requirements of the flange will determine the more suitable process in each case. Note that multi-stage SPIF is the most common alternative to the single-stage 100 Conventional vs. Incremental hole-flanging process to improve the material distribution during a flanging operation. As highlighted by several researchers (Morales-Palma et al.,2017, Skjoedt et al., 2008, 2010, Verbert et al.,2008, Duflou et al.,2010, Liu et al.,2014), multi-stage strategies are able to control the excessive thinning and can improve the dimensional accuracy of the flange. However, apart from an increase in the process complexity, the main drawback is the notable increase in process time as the number of stages (steps) increases. Figure 5.9 Comparison of flange height and thickness for specimens with a 61mm pre cut hole diameter and 8 mm tool radius by SPIF at 0 rpm and press working 5.5. Forming forces and process time Force requirements and production time are determining factors in the selection of a suitable technology. Hole flanging by conventional press working requires heavy machinery and dedicated dies and punches for each flanging size. Assuming that the press is available, the high cost of these accessories renders the process acceptable only large production batches. An incremental hole flanging operation requires considerably less force than a conventional operation, which enables the operation in a standard CNC machine. 101 Analysis of single stage SPIF process applied to the hole flanging operation In this case, the auxiliary structure and forming tools are substantially less expensive and more versatile than conventional hole flanging dies, which renders incremental sheet forming processes profitable for small production batches. However, their main weakness is the high processing time required for a conventional operation. This comparison could be completed by quantifying the level of force and process time that is required for performing the hole flanging tests by single-stage SPIF and conventional press presented in this study. The optimization of these process variables was not the aim of this study; thus, the differences listed in Table 5-3 are estimates of the results expected in industrial operations. Table 5-3 Summary of maximum forming force in z-direction and processing time in both hole flanging operation for tests with a 65 mm pre cut hole diameter and 8 mm tool radius Hole flanging Operation Max. forming force (kN) Processing time (min) Hole milling and deburring Set-up time Hole flanging Single-stage SPIF: feed rate: 1000 mm/min, 0 rpm stepdown: 0.2 mm/rev, 8 mm tool 1.5 15 3 30 Conventional press working: cylindrical punch - 8 mm edge radius punch speed: 0.1 mm/s 30.5 1 1 Table 5-3 summarizes the maximum forming force in the z-direction (vertical) and the processing time required to perform a hole flanging test with a 65 mm pre cut hole diameter and 8 mm tool radius by both technologies. As described in previous chapters, the feed rate and stepdown per revolution used for single-stage SPIF was 1000 mm/min and 0.2 mm/rev, respectively, and the punch speed for press working was 0.1 mm/s. 102 Conventional vs. Incremental hole-flanging Figure 5.10 Evolution of forming force vs. tool travel in zdirection in single-stage incremental and conventional hole flanging for tests of 65 mm pre cut hole diameter and 8 mm tool radius Figure 5.10 depicts the evolution of the tool registered in the z-direction as a function of the tool travel in both forming operations. As expected, the maximum force in conventional forming is one order of magnitude higher than the incremental one. As shown in Table 5-3, the maximum forces are approximately 30.5kN and 1.5kN, that is, a difference of approximately 20 times between operations. Regarding the processing time, the trend reverses. The processing time can be split into three parts: (1) the time for milling and deburring of the initial hole, which is a common operation in both forming processes, (2) the set-up time of the test, that is, the time required to place the specimen, apply the tribological system, and select the process parameters or CNC program, and (3) the hole flanging time, i.e., the time required by the tool to form the flange. Table 5-3 shows the time spent in each of these phases. The flanging time is critical; it is one order of magnitude higher in incremental forming than conventional forming for these tests. This difference can be considerably modified in an industrial context. 103 Analysis of single stage SPIF process applied to the hole flanging operation 5.6. Conclusions Two hole flanging processes conventional press working and the single-stage incremental forming have been critically compared in terms of flangeability, geometrical features of the finished part and required forming force and time. The main conclusions drawn are summarized as follows:  Regarding the strain in the flange, the main difference between both processes is derived from the evolution of the meridional strain. In single-stage SPIF, this strain displays a high positive value around the middle of the flange, which is responsible for a longer and thinner flange thanthe flange in the conventional process.  The analysis based on the FLD satisfactorily reflects the mechanisms that control the flangeability in both processes. In press working, the failure is triggered by necking at the flange edge and controlled by the FLC. In incremental forming, the failure initiates by ductile fracture near the middle of the flange wall and is then controlled by the FFL.  The traditional LFR is not an appropriate parameter for measuring the flangeability in hole flanging by SPIF as it does not capture the physics of failure of this process. A LFR value of approximately 1.6 is obtained for both processes.  Alternatively, the non-dimensional flange height (ℎ   ⁄) and the average thickness ratio ( 3 ⁄) are more suitable for evaluating the flangeability in both processes. In terms of these parameters, the single-stage SPIF exhibits a significant enhancement in formability regarding the press working. A gain in flangeability of approximately 40% and 30% is quantified using ℎ   ⁄ and  3 ⁄, respectively.  Regarding the flange geometry, the single-stage SPIF is able to perform higher and thinner flanges for a given HER parameter. 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