Cutting, Blanking, Shearing & Trimming

Cutting, Blanking, Shearing & Trimming

 

Advanced High-Strength Steels (AHSS) exhibit high degrees of work hardening, resulting in improved forming capabilities compared to conventional HSLA steels. However, the same high work hardening creates higher strength and hardness in sheared or punched edges, leading to reduced edge ductility. Microstructural features in some AHSS grades contribute to their sheared edge performance.  While laser cutting results in less edge damage than mechanical cutting methods, the heat from laser cutting produces a localized hear treatment, changing the strength and hardness at the edge.  Achieving the best formability for chosen processing path requires generating a consistent good quality edge from the cutting operation.

To avoid unexpected problems during a program launch, use production intent tooling as early in the development as possible. This may be a challenge since blanking dies are usually among the last set of tools completed.  In the interim, many companies choose to use laser cut blanks. Tool, blank, and process development must account for the lower-ductility sheared edges in production blanks.

 

Edge Ductility Measurements

This article describes the impact of cutting and cut edge quality on edge ductility.  The primary tests which quantify edge ductility are Hole Expansion Testing, 2-D Edge Tension Testing, and Half Specimen Dome Testing.  These links detail the testing procedures.  The Hole Expansion Testing article has additional information pertaining to the effect of burr orientation and punch shape.

 

Cut Edge Quality

Any mechanical cutting operation such as blanking, piercing, shearing, slitting, or trimming reduces edge ductility.  Each of  these processes generate a zone of high work hardening and a reduced n-value. This work hardened zone can extend one-half metal thickness from the cut edge. This is one reason why edges fail at strains lower than that predicted by the forming limit curve for that particular grade (Note that FLCs were developed based on necking failure, and that edge cracking is a different failure mechanism). 

DP and TRIP steels have islands of martensite located throughout the ferritic microstructure, including at the cut edges. These hard particles act as crack initiators and further reduce the allowable edge stretch. Metallurgical changes to the alloy minimize the hardness differences between the phases, resulting in improved edge ductility.  Laser, EDM or water jet cutting approaches minimize work hardening at the edges and the associated n-value reduction, also leading to improved edge ductility.

Putting shear angles into cutting tools is a well-known approach to reduce cutting forces.  Modifying the cutting tool leads to other benefits in terms of edge ductility. Researchers studied the effects of a beveled punch instead of the traditional flat bottom punch.S-9, S-50, S-52 In these studies, the optimized bevel angle was between 3 and 6 degrees, the shear direction was parallel the rolling direction of the coil with a die clearance of 17%.  With the optimal cutting parameters, the hole expansion ratio increased by 60% when compared to conventional flat punching process.  As expected, a reduction in the maximum shearing force occurred – by more than 50% in certain conditions.  Dropping the shearing force helps reduce the snap through reverse tonnage, leading to longer tool and press life.

Multiple studies examine the trimmed edge quality based on various cutting conditions in mechanical shearing operations and other methods to produce a free edge such as milling and cutting using a laser or water jet. Edge quality varies based on parameters like cutting clearances, shear angles, and rake angles on mechanical shearing operations.

A typical mechanically sheared steel edge has 4 main zones – rollover, burnish, fracture, and burr, as shown in Figure 1.

Figure 1: Cross Section of a Punched Hole Showing the Shear Face Components and Shear Affected Zone S-51

Figure 1: Cross Section of a Punched Hole Showing the Shear Face Components and Shear Affected Zone.K-10

 

Parts stamped from conventional mild and HSLA steels have historically relied on burr height as the main measure of edge quality, where the typical practice targeted a burr height below 10% of metal thickness and slightly larger for thicker steel. Finding a burr exceeded this threshold usually led to sharpening or replacing the trim steels, or less likely, adjusting the clearances to minimize the burr.

Greater burr height is associated with additional cold working and creates stress risers that can lead to edge splitting. These splits, however, are global formability related failures where the steel thins significantly at and around the split, independent of the local formability edge fractures associated with AHSS.  A real-world example is shown in Figure 2, which presents a conventional BH210 steel grade liftgate with an excessive burr in the blank that led to global formability edge splitting in the draw die.  The left image in Figure 2 highlights the burr on the underside of the top blank, with the remainder of the lift below it.  The areas next to the split in the right image of Figure 2 shows the characteristic thinning associated with global formability failures.

Figure 2: Excessive burr on the blank led to a global formability split on the formed liftgate.  The root cause was determined to be dull trim steels resulting in excessive work hardening.U-6

Figure 2: Excessive burr on the blank led to a global formability split on the formed liftgate.  The root cause was determined to be dull trim steels resulting in excessive work hardening.U-6

 

Due to their progressively higher yield and tensile strengths, AHSS grades experience less rollover and smaller burrs. They tend to fracture with little rollover or burr. As such, detailed examination of the actual edge condition under various cutting conditions becomes more significant with AHSS as opposed characterizing edge quality by burr height alone. Examination of sheared edges produced under various trimming conditions, including microhardness testing to evaluate work hardening after cold working the sheared edge, provides insight on methods to improve cut edge formability.  The ideal condition to combat local formability edge fractures for AHSS was to have a clearly defined burnish zone with a uniform transition to the fracture zone. The fracture zone should also be smooth with no voids, secondary shear or edge damage (Figure 3).

Figure 3: Ideal sheared edge with a distinct burnish zone and a smooth fracture zone (left) and a cross section of the same edge (right).U-6

Figure 3: Ideal sheared edge with a distinct burnish zone and a smooth fracture zone (left) and a cross section of the same edge (right).U-6

  

If clearances are too small, secondary shear can occur and the potential for voids due to the multiphase microstructure increases, as indicated in Figure 4.  Clearances that are too large create additional problems that include excessive burrs and voids. A nonuniform transition from the burnish zone to the fracture zone is also undesirable. These non-ideal conditions create propagation sites for edge fractures. 

Figure 4: Sheared edge with the trim steel clearance too small (left) and a cross section of the same edge (right) showing a micro crack on the edge. Tight clearance leading to secondary shear increases the likelihood of edge fracture.U-6

Figure 4: Sheared edge with the trim steel clearance too small (left) and a cross section of the same edge (right) showing a micro crack on the edge. Tight clearance leading to secondary shear increases likelihood of edge fracture.U-6

 

There are multiple causes for a poor sheared edge condition, including but not limited to:

  • the die clearance being too large or too small, 
  • a cutting angle that is too small, 
  • worn, chipped, or damaged tooling,
  • improperly ground or sharpened tooling,
  • improper die material, 
  • improperly heat-treated die material, 
  • improper (or non-existent) coating on the tooling, 
  • misaligned die sections, 
  • worn wear plates, and
  • out of level presses or slitting equipment. 

The higher loads required to shear AHSS with increasingly higher tensile strength creates additional deflection of dies and processing equipment. This deflection may alter clearances measured under a static condition once the die, press, or slitting equipment is placed under load. As a large percentage of presses, levelers, straighteners, blankers, and slitting equipment were designed years ago, the significantly higher loads required to process today’s AHSS may exceed equipment beyond their design limits, dramatically altering their performance.

A rocker panel formed from DP980 provides a good example showing the influence of cut edge quality. A master coil was slit into several narrower coils (mults) before being shipped to the stamper.  Only a few mults experienced edge fractures, which all occurred along the slit edge. Understanding that edge condition is critical with respect to multiphase AHSS, the edge condition of the “good” mults and the “bad” mults were examined under magnification. The slit edge from a problem-free lift (Figure 5) has a uniform burnish zone with a uniform transition to the smooth fracture zone. This is in contrast with Figure 6, from the slit edge from a different mult of the same coil in which every blank fractured at the slit edge during forming. This edge exhibits secondary shear as well as a thick burnish zone with a non-uniform transition from the burnish zone to the fracture zone.

Figure 5: Slit edges on a lift of blanks that successfully produced DP980 rocker panels. Note the uniform transition from the burnish zone to the fracture zone with a smooth fracture zone as well.U-6

Figure 5: Slit edges on a lift of blanks that successfully produced DP980 rocker panels. Note the uniform transition from the burnish zone to the fracture zone with a smooth fracture zone as well.U-6

 

Figure 6: Slit edges on a lift of blanks from the same master coil that experienced edge fractures during forming. Note the obvious secondary shear as well as the thicker, nonuniform transition from the burnish to the fracture zone.U-6

Figure 6: Slit edges on a lift of blanks from the same master coil that experienced edge fractures during forming. Note the obvious secondary shear as well as the thicker, nonuniform transition from the burnish to the fracture zone.U-6

 

Cutting Clearances: Burr Height and Tool Wear

Cutting and punching clearances should be increased with increasing sheet material strength. The clearances range from about 6% of the sheet material thickness for mild steel up to 16% or even higher as the sheet metal tensile strength exceeds 1400 MPa.

A study C-2  compared the tool wear and burr height formation associated with punching mild steel and several AHSS grades. In addition to 1.0 mm mild steel (140 MPa yield strength, 270 MPa tensile strength, 38% A80 elongation), AHSS grades tested were 1.0 mm samples of DP 350Y600T (A80=20%), DP 500Y800T (A80=8%), and MS 1150Y1400T (A80 = 3%).  Tests of mild steel used a 6% clearance and W.Nr. 1.2363 / AISI A2 tool steel hardened to 61 HRC.  The AHSS tests used engineered tool steels made from powder metallurgy hardened to 60-62 HRC.  The DP 350/600 tests were run with a TiC CVD coating, and a 6% clearance. Tool clearances were 10% for the MS 1150Y1400T grade and 14% for DP 500Y800T.

In the Tool Wear comparison, the cross-section of the worn punch was measured after 200,000 hits.  Punches used with mild steel lost about 2000 μm2 after 200,000 hits, and is shown in Figure 7 normalized to 1. The relative tool wear of the other AHSS grades are also shown, indicating that using surface treated high quality tool steels results in the same level of wear associated with mild steels punched with conventional tools.

Figure 7: Tool wear associated with punching up to DP 500Y800T using surface treated high quality tool steels is comparable to mild steel punched with conventional tools. C-2

Figure 7: Tool wear associated with punching up to DP 500Y800T using surface treated high quality tool steels is comparable to mild steel punched with conventional tools.C-2

 

Figure 8 shows the burr height test results, which compared burr height from tests using mild steel punched with conventional tool steel and two AHSS grades (DP 500Y800T and MS 1150Y1400T) punched with a PM tool steel. The measured burr height from all AHSS and clearance combinations evaluated were sufficiently similar that they are shown as a single curve.

Figure 8  Burr height comparison for mild steel and two AHSS grades as a function of the number of hits. Results for DP 500Y800T and Mart 1150Y1400T are identical and shown as the AHSS curve.C-2

Figure 8:  Burr height comparison for mild steel and two AHSS grades as a function of the number of hits. Results for DP 500Y800T and Mart 1150Y1400T are identical and shown as the AHSS curve.C-2

 

Testing of mild steel resulted in the expected performance where burr height increases continuously with tool wear and clearance, making burr height a reasonable indicator of when to sharpen punching or cutting tools.  However, for the AHSS grades studied, burr height did not increase with more hits. It is possible that the relatively lower ductility AHSS grades are not capable of reaching greater burr height due to fracturing, where the more formable mild steel continues to generate ever-increasing burr height with more hits and increasing tool wear.

Punching AHSS grades may require a higher-grade tool steel, possibly with a surface treatment, to avoid tool wear, but tool regrinding because of burrs may be less of a problem.  With AHSS, engineered tool steels may provide longer intervals between sharpening, but increasing burr height alone should not be the only criterion to initiate sharpening: cut edge quality as shown in the above figures appears to be a better indicator.  Note that regrinding a surface treated tool steel removes the surface treatment. Be sure to re-treat the tool to achieve targeted performance.

 

Cutting Clearances: General Recommendations

Depending on the source, the recommended die clearance when shearing mild steels is 5% to 10% of metal thickness. For punched holes, these represent per-side values.  Although this may have been satisfactory for mild steels, the clearance should increase as the tensile strength of the sheet metal increases.  

The choice of clearance impacts other aspects of the cutting process.  Small cutting clearances require improved press and die alignment, greater punching forces, and cause greater punch wear from abrasion. As clearance increases, tool wear decreases, but rollover on the cut edge face increases, which in the extreme may lead to a tensile fracture in the rollover zone (Figure 9). Also, a large die clearance when punching high strength materials with a small difference in yield and tensile strength (like martensitic grades) may generate high bending stresses on the punch edge, which increases the risk of chipping.

Figure 9: Large rollover may lead to tensile fracture in the rollover zone.

Figure 9: Large rollover may lead to tensile fracture in the rollover zone.

 

Figure 10 compares cut edge appearance after punching a martensitic steel with 1400 MPa tensile strength using either 6% or 14% clearance.  The larger clearance is associated with greater rollover, but a cleaner cut face.

Figure 10: Cut edge appearance after punching CR 1400T-MS with 6% (left) and 14% (right) die clearance. The bottom images show the edge appearance for the full sheet thickness,  Note using 6% clearance resulted in minimal rollover, but uneven burnish and fracture surfaces.  In contrast, 14% clearance led to noticeable rollover, but a clean burnish and fracture surface. T-20

Figure 10: Cut edge appearance after punching CR 1400T-MS with 6% (left) and 14% (right) die clearance. The bottom images show the edge appearance for the full sheet thickness,  Note using 6% clearance resulted in minimal rollover, but uneven burnish and fracture surfaces.  In contrast, 14% clearance led to noticeable rollover, but a clean burnish and fracture surface.T-20

 

A comparison of the edges of a 2 mm thick complex phase steel with 700 MPa minimum tensile strength produced under different cutting conditions is presented in Figure 11. The left image suggests that either the cutting clearance and/or the shearing angle was too large. The right image shows an optimal edge likely to result in good edge ductility.

Figure 11: Cut edge appearance of 2mm HR 700Y-MC, a complex phase steel. The edge on the right is more likely to result in good edge ductility.T-20

Figure 11: Cut edge appearance of 2 mm HR 700Y-MC, a complex phase steel. The edge on the right is more likely to result in good edge ductility.T-20

 

The recommended clearance is a function of the sheet grade, thickness, and tensile strength.  Figures 12 to 15 represent general recommendations from several sources.

Figure 12:  Recommended Clearance as a Function of Grade and Sheet Thickness. T-23

Figure 12:  Recommended Clearance as a Function of Grade and Sheet Thickness.T-23

 

Figure 13: Recommended Cutting Clearance for Punching.D-15

Figure 13: Recommended Cutting Clearance for Punching.D-15

 

Figure 14: Recommended die clearance for blanking/punching advanced high strength steel. T-20

Figure 14: Recommended die clearance for blanking/punching advanced high strength steel.T-20

 

Figure 15:  Multiply the clearance on the left with the scaling factor in the right to reach the recommended die clearance.D-16

Figure 15:  Multiply the clearance on the left with the scaling factor in the right to reach the recommended die clearance.D-16

 

Figure 16 highlights the effect of cutting clearance on CP1200, and reinforces that the historical rule-of-thumb guidance of 10% clearance does not apply for all grades. In this studyU-3, increasing the clearance from 10% to 15% led to a significant improvement in hole expansion. The HER resulting from a 20% clearance was substantially better than that from a 10% clearance, but not as good as achieved with a 15% clearance. These differences will not be captured when testing only to the requirements of ISO 16630, which specifies the use of 12% clearance.

Figure 16: Effect of hole punching clearance on hole expansion of Complex Phase steel grade CP1200.U-3

Figure 16: Effect of hole punching clearance on hole expansion of Complex Phase steel grade CP1200.U-3

 

Cutting speed influences the cut edge quality, so it also influences the optimal clearance for a given grade. In a study published in 2020G-49, higher speeds resulted in better sheared edge ductility for all parameters evaluated, with those edges having minimal rollover height, smoother sheared surface and negligible burr. Two grades were evaluated: a dual phase steel with 780MPa minimum tensile strength and a 3rd Generation steel with 980 MPa minimum tensile strength.

Metallurgical characteristics of the sheet steel grade also affects hole expansion capabilities. Figure 17 compares the HER of DP780 from six global suppliers. Of course, the machined edge shows the highest HER due to the minimally work-hardened edge. Holes formed with 13% clearance produced greater hole expansion ratios than those formed with 20% clearance, but the magnitude of the improvement was not consistent between the different suppliers.K-56

Figure 17: Cutting clearance affects hole expansion performance in DP780 from 6 global suppliers Citation K-56

Figure 17: Cutting clearance affects hole expansion performance in DP780 from six global suppliers.K-56

 

 

Punch Face Design

Practitioners in the field typically do not cut perpendicular to the sheet surface – angled punches and blades are known to reduce cutting forces.  For example, long shear blades might have a 2 to 3 degree angle on them to minimize peak tonnages.  There are additional benefits to altering the punch profile and impacting angle.

Snap-though or reverse tonnage results in stresses which may damage tooling, dies, and presses. Tools may crack from fatigue.  Perhaps counter to conventional thinking, use of a coated punch increases blanking and punching forces. The coating leads to lower friction between the punch and the sheet surface, which makes crack initiation more difficult without using higher forces. 

Unlike a coated tool, a chamfered punch surface reduces blanking and punching forces.  Figure 18 compares the forces to punch a 5 mm diameter hole in 1 mm thick MS-1400T using different punch shapes. A chamfered punch was the most effective in reducing both the punching force requirements and the snap-through tonnage (the shock waves and negative tonnage readings in Figure 18).  The chamfer should be large enough to initiate the cut before the entire punch face is in contact with the sheet surface.  A larger chamfer increases the risk of plastic deformation of the punch tip.T-20

Figure 16: A chamfered punch reduces peak loads and snap-through tonnage.K-15

Figure 18: A chamfered punch reduces peak loads and snap-through tonnage.K-15

 

A different study P-16 showed more dramatic benefits. Use of a rooftop punch resulted in up to an 80% reduction in punching force requirements compared with a flat punch, with a significant reduction in snap-through tonnage.  Cutting clearance had only minimal effect on the results. (Figure 19)

Figure 17: A rooftop-shaped punch leads to dramatic reductions in punch load requirements and snap-through tonnage.P-16

Figure 19: A rooftop-shaped punch leads to dramatic reductions in punch load requirements and snap-through tonnage.P-16

 

Use of a beveled punch (Figure 20) provides similar benefits.  A study S-52 comparing DP 500/780 and DP 550/980 showed a reduction in the maximum piercing force of more than 50% with the use of a beveling angle between 3 and 6 degrees. The shearing force depends also upon the die clearance during punching, with the optimum performance seen with 17% die clearance. The optimal punching condition results in more than 60% improvement in the hole expansion ratio when compared to conventional flat head punching process.  The optimal bevel cut edge in Figure 21 shows a uniform burnish zone with a uniform transition to the smooth fracture zone – the known conditions to produce a high-ductility edge.

Figure 18: Schematic showing a beveled punch S-52

Figure 20: Schematic showing a beveled punch.S-52

 

Figure 19: A bevel cut edge showing uniform burnish zone with a uniform transition to the smooth fracture zone.S-52

Figure 21: A bevel cut edge showing uniform burnish zone with a uniform transition to the smooth fracture zone.S-52

 

Effect of Edge Preparation Method on Ductility

A flat trim condition where the upper blade and lower blade motions are parallel and there is no shear rake angle is known to produce a trimmed edge with limited edge stretchability (Figure 22, left image).  In addition to split parts, tooling damage and unexpected down time results.  Metal stampers have known that shearing with a rake angle Figure 22, right image) will reduce cutting forces compared with using a flat cut.  With advanced high strength steels, there is an accompanying reduction in forming energy requirements of up to 20% depending on the conditions, which represents a tremendous drop in snap-through or reverse tonnage.  Figure 22 visually describes the upper and lower blade rake angles and the shear rake angle.

Figure 20: Flat trim (left) and shear trim (right) conditions showing rake angle definitions.S-53

Figure 22: Flat trim (left) and shear trim (right) conditions showing rake angle definitions.S-53

   

Researchers have also found that it is possible to increase sheared edge ductility with optimized rake angles. Citation S-53 used 2-D Edge Tension Testing and the Half-Specimen Dome Test to qualify the effects of these rake angles, and determine the optimum settings.  After preparing the trimmed edge with the targeted conditions, the samples were pulled in a tensile test or deformed using a hemispherical punch. The effect of the trimming conditions was seen in the measured elongation values and the strain at failure, respectively.  The results are summarized in Figures 23-25.  Some of the tests also evaluated milled, laser trimmed, and water jet cut samples. Shear Trim 1, 2, and 3 refer to the shear trim angle in degrees. The optimized shear condition also includes a 6-degree rake angle on both the upper and lower blades, as defined in Figure 22.  

Conclusions from this study include:

  • Mechanically shearing the edge cold works the steel and reduces the work hardening exponent (n-value), leading to less edge stretchability. 
  • Samples prepared with processes that avoided cold working the edges, like laser or water jet cutting outperformed mechanically sheared edges.  
  • Optimizing the trim shear conditions or polishing a flat trimmed edge approaches what can be achieved with laser trimming and water jet cutting.
  • Shearing parameters such as clearance, shear angle and rake angle also play a large part in improving edge stretch. 
Figure 21: Effect of edge preparation on stretchability as determined using a tensile test for DP350Y600T (left) and DP550Y980T (right).S-53

Figure 23: Effect of edge preparation on stretchability as determined using a tensile test for DP 350Y600T (left) and DP 550Y980T (right).S-53

 

Figure 22: Effect of edge preparation on stretchability as determined using a dome test for DP350Y600T (left) and DP550Y980T (right).S-53

Figure 24: Effect of edge preparation on stretchability as determined using a dome test for DP 350Y600T (left) and DP 550Y980T (right).S-53

 

Figure 23: Optimizing the trim shear conditions or polishing a flat trimmed edge approaches what is achievable with laser trimming and water jet cutting. Data from dome testing of DP 350Y/600T.S-53

Figure 25: Optimizing the trim shear conditions or polishing a flat trimmed edge approaches what is achievable with laser trimming and water jet cutting. Data from dome testing of DP 350Y/600T.S-53

 

The optimal edge will have no mechanical damage and no microstructural changes as you go further from the edge.  Any process that changes the edge quality from the bulk material can influence performance.  This includes the mechanical damage from shearing operations, which cold works the edge leading to a reduction in ductility.  Laser cutting also changes the edge microstructure, since the associated heat input is sufficient to alter the engineered balance of phases which give AHSS grades their unique properties.  However, the heat from laser cutting is sometimes advantageous, such as in the creation of locally softened zones to improve cut edge ductility in some applications of press hardening steels.

The effects of edge preparation on the shear affected zone is presented in Figure 26.  A flatter profile of the Vickers microhardness reading measured from the as-produced edge into the material indicates the least work-hardening and mechanical damage resulting from the edge preparation method, and therefore should result in the greatest edge ductility.  This is certainly the case for water jet cutting, where a flat hardness profile in Figure 26 correlates with the highest ductility measurements in Figures 22 to 25. Unfortunately, water jet cutting is not always practical, and introduces the risk of rust forming at the newly cut edge.

Figure 24:  Microhardness profile starting at cut edge generated using different methods.  Left image is from S-53, and right image is from C-13

Figure 26:  Microhardness profile starting at cut edge generated using different methods.  Left image is from Citation S-53, and right image is from C-13.

 

Two-stage piercing is another method to reduce edge strain hardening effects. Here, a conventional piercing operation is followed by a shaving operation which removes the work-hardened material created in the first step, as illustrated in Figure 27.P-17 A related studyF-10 evaluated this method with a 4 mm thick complex phase steel with 800 MPa tensile strength.  Using the configuration documented in this reference, single-stage shearing resulted in a hole expansion ratio of only 5%, where the addition of the shaving operation improved the hole expansion ratio to 40%.

Figure 25: Two-stage piercing improves cut edge ductility. Image adapted from P-17

Figure 27: Two-stage piercing improves cut edge ductility. Image adapted from Citation P-17.

 

Figure 28 highlights the benefits of two-stage pre-piercing for specific grades, showing a 2x to 4x improvement in hole expansion ratio for the grades presented.

Figure 27: Pre-piercing improves the hole expansion ratio of AHSS Grades.S-10

Figure 28: Pre-piercing improves the hole expansion ratio of AHSS Grades.S-10

Key Points

  • Clearances for punching, blanking, and shearing should increase as the strength of the material increases, but only up to a point. At the highest strengths, reducing clearance improves tool chipping risk.
  • Lower punch/die clearances lead to accelerated tool wear. Higher punch/die clearances generate more rollover/burr.
  • ISO 16630, the global specification for hole expansion testing, specifies the use of 12% punch-to-die clearance. Optimized clearance varies by grade, so additional testing may prove insightful.
  • Recommended clearance as a percentage of sheet thickness increases with thickness, even at the same strength level. 
  • Burr height increases with tool wear and increasing die clearances for shearing mild steel, but AHSS tends to maintain a constant burr height. This means extended intervals between tool sharpening may be possible with AHSS parts, providing edge quality and edge performance remain acceptable.
  • Edge preparation methods like milling, laser trimming, and water-jet cutting minimize cold working at the edges, resulting in the greatest edge ductility,
  • Laser cut blanks used during early tool tryout may not represent normal blanking, shearing, and punching quality, resulting in edge ductility that will not occur in production.  Using production-intent tooling as early as possible in the development stage minimizes this risk.
  • Shear or bevel on punches and trim steel reduces punch forces, minimizes snap-through reverse tonnage, and improves edge ductility.
  • Mild steel punched with conventional tools and AHSS grades punched with surface treated engineered PM tool steels experience comparable wear.
  • Maintenance of key process variables, such as clearance and tool condition, is critical to achieving long-term edge stretchability. 
  • The optimal edge appearance shows a uniform burnish zone with a uniform transition to a smooth fracture zone.

 

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Global vs Local Formability

Global vs Local Formability

Most sheet metals have different “Global” and “Local” forming capabilities, so it is critical to understand their meaning to optimize grade selection, processing, and usage.

Historically, most fabricators needed to consider only global formability when designing and stamping parts. Drawing, plane-strain tension, and stretch forming are “global” forming modes where deformation occurs in the plane of the sheet over relatively large regions of material. Tensile failures or necking failures, where the steel progressively thins during forming, are a characteristic of global formability failures. The strains across a stamped part start low and evenly distributed but begin to concentrate as the punch approaches bottom-dead-center. Critical thinning (also called necking) occurs if the strains induced by the part shape and forming process exceed the forming limit of the chosen sheet metal. Forming simulation software packages have reliably shown the ability to accurately predict global formability concerns and hot spots with the use of inputs like tensile test data and the correct forming limit curve.

Local formability failure modes are an entirely different failure condition, where fractures occur out of the plane of the sheet in response to concentrated deformation created when forming localized features like stretch flanges, extruded holes, or bends around a radius too small for the selected steel grade. These failures typically occur without any observable thinning or necking (Figure 1). Forming simulation software that considers only the forming limit curve or maximum thinning as the failure criteria cannot predict local formability failures.

Figure 1: A fracture related to insufficient local formability. Note the lack of thinning near the fracture.

Figure 1: A fracture related to insufficient local formability. Note the lack of thinning near the fracture. H-5

 

Cutting conditions and edge stresses developed in blanking and slitting operations play a significant role in limiting local formability. Trim steel clearances, shear angles, trim steel materials, tool sharpness and design, steel rolling direction, and part design considerations are all important (Figure 2).

Figure 2: Cut edge uniformity influences edge quality.

Figure 2: Cut edge uniformity influences edge quality.U-6

 

The absence of global thinning leading to the onset of localized necking highlights the importance of clearly defining the actual mode of failure before trying to identify possible solutions. Examination of the edge of the part where the failure is occurring is the best way to accomplish this. Figure 3 shows a photo of a typical local formability related edge fracture where no observable thinning/necking before fracture occurs.

Figure 3: A typical local formability edge fracture viewed from different angles, with no appreciable thinning prior to fracture.U-6

Figure 3: A typical local formability edge fracture viewed from different angles, with no appreciable thinning prior to fracture.U-6

 

Some AHSS grades are more prone to these local formability failures. A reduced hardness difference between microstructural phases appears to improve local formability. At a given tensile strength level, there may be grades with improved global formability, improved local formability, or a balanced approach available. Choosing the optimum grade requires understanding of the manufacturing die process and the functionality requirements associated with the part and its application.

Figure 4 shows an edge fracture on a DP980 rocker panel. This panel experienced edge fractures during production, occurring in embossments along the edge of the part, with no evidence of necking at the fracture site. This is a classic local formability failure, reinforcing the need to inspect the conditions at the fracture zone of AHSS to determine whether the root failure is global or local formability. Process and die solutions differ based on the mode of failure.

Figure 4: Classic local formability related edge fracture on an embossment within a DP980 part.U-6

Figure 4: Classic local formability related edge fracture on an embossment within a DP980 part.U-6

 

Grade dependency is highlighted in Figure 5, which compares edges of HSLA 50SK on the left and DP500Y/780T on the right after stretch-bend testing under conditions to produce fracture. The HSLA sample shows the characteristic thinning down associated with necking. The DP steel did not have a visible neck at the failure location.S-11

Figure 5:  Left: HSLA 50SK, showing thinning at the fracture location, which is typical for global formability failures Right: DP 500Y/780T, showing no thinning at the fracture location, which is typical for local formability failures.S-11

Figure 5:  Left: HSLA 50SK, showing thinning at the fracture location, which is typical for global formability failures.  Right: DP 500Y/780T, showing no thinning at the fracture location, which is typical for local formability failures.S-11

 

Mechanical tests now being employed to better quantify and characterize the unique local formability related failure modes associated with AHSS include:

Summary of Global Formability

  • The resistance to localized thinning (necking) is the key to global formability.
  • Stamped parts get progressively thinner as the press stroke approaches bottom-dead-center. Rapid thinning in critical areas lead to the onset of localized necking. Global formability promotes the ability to reach deeper draw depths before initiation of necking.
  • Measures of global formability include the work-hardening exponent (n-value), the uniform elongation value, and the total elongation value determined in a tensile test, along with the forming limit curve (FLC).
  • Global formability failures occur when a through-thickness volume of the formed sheet steel exceeds this forming limit and begins to neck.
    Necking failure typically should not occur if the global formability limit is exceeded on only the outer surface of steel bent over a radius or expanded at a cut edge and not through the full thickness of the metal.
  • Uniform elongation measures the resistance to local necking. The global formability limit corresponds to Uniform Elongation in the strain state achieved in a tensile test sample. The Forming Limit Curve (FLC) encompasses all strain states.
  • For mild and conventional high strength steels, there is a large difference between the necking limit and the ultimate fracture limit. This corresponds to the difference between Uniform Elongation and Total Elongation in a tensile test, where the value of total elongation is typically twice that of uniform elongation.

 

Summary of Local Formability

  • The resistance to fracture is the key to local formability.
  • Local formability promotes fracture resistance in response to creating local product features like stretch flanges, extruded holes, and tight-radius bends.
  • Local formability tests which evaluate the stretchability of appropriately prepared edges by deforming them with a punch may have limited reproducibility between labs due to the influence of different sample preparation methods. Determining parameters like True Fracture Strain, Reduction in Area at Fracture, or Thickness Strain at Fracture with a well-defined tensile test typically results in repeatable and reproducible values.
  • Local formability failure occurs when strains at the surface or edge exceed the ultimate fracture limit, a value that is higher than the necking limit determined by the Forming Limit Curve.
  • Local formability failures occur more frequently on higher strength steels partly because of a smaller difference between the necking limit determined by the Forming Limit Curve and ultimate fracture limit.
  • Work hardening and damage from edge preparation methods like shearing further reduce edge formability to values typically lower than indicated by the Forming Limit Curve, especially for AHSS. The strain localization at the interface between hard and soft phases in AHSS also contribute to an increased risk of local formability failures.
  • The lack of a localized neck is a characteristic trait associated with of local formability failures.
Cutting, Blanking, Shearing & Trimming

N-Value

N-Value, The Strain Hardening Exponent

Metals get stronger with deformation through a process known as strain hardening or work hardening, resulting in the characteristic parabolic shape of a stress-strain curve between the yield strength at the start of plastic deformation and the tensile strength.

Work hardening has both advantages and disadvantages. The additional work hardening in areas of greater deformation reduces the formation of localized strain gradients, shown in Figure 1.

Figure 1: Higher n-value reduces strain gradients, allowing for more complex stampings. Lower n-value concentrates strains, leading to early failure.

Figure 1: Higher n-value reduces strain gradients, allowing for more complex stampings. Lower n-value concentrates strains, leading to early failure.

 

Consider a die design where deformation increased in one zone relative to the remainder of the stamping. Without work hardening, this deformation zone would become thinner as the metal stretches to create more surface area. This thinning increases the local surface stress to cause more thinning until the metal reaches its forming limit. With work hardening the reverse occurs. The metal becomes stronger in the higher deformation zone and reduces the tendency for localized thinning. The surface deformation becomes more uniformly distributed.

Although the yield strength, tensile strength, yield/tensile ratio and percent elongation are helpful when assessing sheet metal formability, for most steels it is the n-value along with steel thickness that determines the position of the forming limit curve (FLC) on the forming limit diagram (FLD). The n-value, therefore, is the mechanical property that one should always analyze when global formability concerns exist. That is also why the n-value is one of the key material related inputs used in virtual forming simulations.

Work hardening of sheet steels is commonly determined through the Holloman power law equation:

where
σ is the true flow stress (the strength at the current level of strain),
K is a constant known as the Strength Coefficient, defined as the true strength at a true strain of 1,
ε is the applied strain in true strain units, and
n is the work hardening exponent

Rearranging this equation with some knowledge of advanced algebra shows that n-value is mathematically defined as the slope of the true stress – true strain curve. This calculated slope – and therefore the n-value – is affected by the strain range over which it is calculated. Typically, the selected range starts at 10% elongation at the lower end to the lesser of uniform elongation or 20% elongation as the upper end. This approach works well when n-value does not change with deformation, which is the case with mild steels and conventional high strength steels.

Conversely, many Advanced High-Strength Steel (AHSS) grades have n-values that change as a function of applied strain. For example, Figure 2 compares the instantaneous n-value of DP 350/600 and TRIP 350/600 against a conventional HSLA350/450 grade. The DP steel has a higher n-value at lower strain levels, then drops to a range similar to the conventional HSLA grade after about 7% to 8% strain. The actual strain gradient on parts produced from these two steels will be different due to this initial higher work hardening rate of the dual phase steel: higher n-value minimizes strain localization.

Figure 2: Instantaneous n-values versus strain for DP 350/600, TRIP 350/600 and HSLA 350/450 steels.K-1

Figure 2: Instantaneous n-values versus strain for DP 350/600, TRIP 350/600 and HSLA 350/450 steels.K-1

 

As a result of this unique characteristic of certain AHSS grades with respect to n-value, many steel specifications for these grades have two n-value requirements: the conventional minimum n-value determined from 10% strain to the end of uniform elongation, and a second requirement of greater n-value determined using a 4% to 6% strain range.

Plots of n-value against strain define instantaneous n-values, and are helpful in characterizing the stretchability of these newer steels. Work hardening also plays an important role in determining the amount of total stretchability as measured by various deformation limits like Forming Limit Curves.

Higher n-value at lower strains is a characteristic of Dual Phase (DP) steels and TRIP steels. DP steels exhibit the greatest initial work hardening rate at strains below 8%. Whereas DP steels perform well under global formability conditions, TRIP steels offer additional advantages derived from a unique, multiphase microstructure that also adds retained austenite and bainite to the DP microstructure. During deformation, the retained austenite is transformed into martensite which increases strength through the TRIP effect.  This transformation continues with additional deformation as long as there is sufficient retained austenite, allowing TRIP steel to maintain very high n-value of 0.23 to 0.25 throughout the entire deformation process (Figure 2). This characteristic allows for the forming of more complex geometries, potentially at reduced thickness achieving mass reduction. After the part is formed, additional retained austenite remaining in the microstructure can subsequently transform into martensite in the event of a crash, making TRIP steels a good candidate for parts in crush zones on a vehicle.

Necking failure is related to global formability limitations, where the n-value plays an important role in the amount of allowable deformation at failure. Mild steels and conventional higher strength steels, such as HSLA grades, have an n-value which stays relatively constant with deformation. The n-value is strongly related to the yield strength of the conventional steels (Figure 3).

Figure 3: Experimental relationship between n-value and engineering yield stress for a wide range of mild and conventional HSS types and grades.K-2

Figure 3: Experimental relationship between n-value and engineering yield stress for a wide range of mild and conventional HSS types and grades.K-2

 

N-value influences two specific modes of stretch forming:

  1. Increasing n-value suppresses the highly localized deformation found in strain gradients (Figure 1).
    A stress concentration created by character lines, embossments, or other small features can trigger a strain gradient. Usually formed in the plane strain mode, the major (peak) strain can climb rapidly as the thickness of the steel within the gradient becomes thinner. This peak strain can increase more rapidly than the general deformation in the stamping, causing failure early in the press stroke. Prior to failure, the gradient has increased sensitivity to variations in process inputs. The change in peak strains causes variations in elastic stresses, which can cause dimensional variations in the stamping. The corresponding thinning at the gradient site can reduce corrosion life, fatigue life, crash management and stiffness.  As the gradient begins to form, low n-value metal within the gradient undergoes less work hardening, accelerating the peak strain growth within the gradient – leading to early failure. In contrast, higher n-values create greater work hardening, thereby keeping the peak strain low and well below the forming limits. This allows the stamping to reach completion.
  2. The n-value determines the allowable biaxial stretch within the stamping as defined by the forming limit curve (FLC).
    The traditional n-value measurements over the strain range of 10% – 20% would show no difference between the DP 350/600 and HSLA 350/450 steels in Figure 2. The approximately constant n-value plateau extending beyond the 10% strain range provides the terminal or high strain n-value of approximately 0.17. This terminal n-value is a significant input in determining the maximum allowable strain in stretching as defined by the forming limit curve. Experimental FLC curves (Figure 4) for the two steels show this overlap.

    Figure 4: Experimentally determined Forming Limit Curves for mils steel, HSLA 350/450, and DP 350/600, each with a thickness of 1.2mm.K-1

    Figure 4: Experimentally determined Forming Limit Curves for mils steel, HSLA 350/450, and DP 350/600, each with a thickness of 1.2mm.K-1

Whereas the terminal n-value for DP 350/600 and HSLA 350/450 are both around 0.17, the terminal n-value for TRIP 350/600 is approximately 0.23 – which is comparable to values for deep drawing steels (DDS). This is not to say that TRIP steels and DDS grades necessarily have similar Forming Limit Curves. The terminal n-value of TRIP grades depends strongly on the different chemistries and processing routes used by different steelmakers. In addition, the terminal n-value is a function of the strain history of the stamping that influences the transformation of retained austenite to martensite. Since different locations in a stamping follow different strain paths with varying amounts of deformation, the terminal n-value for TRIP steel could vary with both part design and location within the part. The modified microstructures of the AHSS allow different property relationships to tailor each steel type and grade to specific application needs.

Methods to calculate n-value are described in Citations A-43, I-14, J-13.