![Cutting, Blanking, Shearing & Trimming]()
Cutting, 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.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
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
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 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 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 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
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 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
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 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 13: Recommended Cutting Clearance for Punching.D-15
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 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
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 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 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 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 20: Schematic showing a beveled punch.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 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 23: Effect of edge preparation on stretchability as determined using a tensile test for DP 350Y600T (left) and DP 550Y980T (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 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 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 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 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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Press Hardened Steels
Overview
The initial press hardening steels of the 1970s were delivered bare, without a galvanized or aluminized layer for corrosion protection (i.e., uncoated). During the heating process, an oxide layer of FeOx forms if the furnace atmosphere is not controlled. Through the years, several coating technologies have been developed to solve the following problems of uncoated steelsF-14, F-33:
- Scale formation, which causes abrasive wear and requires a secondary shotblasting process before welding,
- Decarburization, which leads to softening close to the surface,
- Risk of corrosion.
The first commercially available coating on press hardening steels was patented in 1998. The coating was designed to solve the scaling problem, but it also offered some corrosion resistance.C-24 Since the coating composition is primarily aluminium, with approximately 9% silicon, it is usually referred to as AlSi, Al-Si, or AS.
Coating thickness is nominally 25 μm (75 g/m2) on each side and referenced as AS150. A more recent offering is a thinner coating of 13 μm (AS80).A-51 The AS coating requires a special heating curve and soaking time for better weldability, corrosion resistance and running health of the furnace. Most OEMs now include furnace dew point limitations to reduce/avoid hydrogen embrittlement risk.
In 2005, Volkswagen was looking for a method to manufacture deep drawn transmission tunnels and other complex-to-form underbody components using press hardened steels. Although AS coatings were available, parts could not be formed to the full draw depth using the direct process, and AS coated blanks cracked during the cold forming portion of the two-step hybrid process. Using uncoated blanks led to severe scale formation, which increased the friction coefficient in hot forming. For this particular problem, a varnish coating was developed. The coating was applied at a steel mill, and shipped to Volkswagen’s stamping plant. The parts were first cold pre-formed and then heated in a furnace, as seen in Figure 1a. Hot pre-forms were then deep drawn to tunnels. As shown in Figure 1b, scale formed on parts which did not have the coating. A varnish coated blank could be cold formed without any scale, Figure 1c.S-63, F-34 Since then, some other varnish coatings also have been developed.
![Figure 1: Transmission tunnel of 2005 Volkswagen Passat: (a) hot forming of pre-form, and final parts: (a) uncoated blank would suffer from scaling, (c) scale-free parts can be formed from varnish-coated blanks [REFERENCE 7]](https://ahssinsights.org/wp-content/uploads/2021/02/Fig1_Transmission-tunnel-of-2005-Volkswagen-Passat.jpg)
Figure 1: Transmission tunnel of 2005 Volkswagen Passat: (a) hot forming of pre-form, and final parts: (a) uncoated blank would suffer from scaling, (c) scale-free parts can be formed from varnish-coated blanks.F-34
In car bodies, components that are sealed from external moisture are referred to as dry areas. These areas have low risk of corrosion. Areas that may be exposed to moisture are wet areas. Precautions must be taken to avoid corrosion of the sheet metal, such as using galvanized or pre-coated steel. Sealants can also be applied to joints to keep out moisture. The presence of humidity in these areas increases the risk of forming a galvanic cell, leading to accelerated corrosion. These areas have higher risk of corrosion and may require additional measures. Figure 2 shows dry and wet areas. In this figure, parts colored with yellow may be classified as wet or dry, depending on the vehicle design and the OEMs requirements.G-41
![Figure 2: Dry and wet areas in a car body. [REFERENCE 8]](https://ahssinsights.org/wp-content/uploads/2021/02/Fig2_Dry-and-wet-areas-in-a-car-body.jpg)
Figure 2: Dry and wet areas in a car body.G-41
An estimated ~40% of press hardened components are in dry areas. Thus, high corrosion protection is desired in the 60% of all press hardened components which are employed in wet areas.B-48 Zn-based coatings are favored for their cathodic protection, but require tight process control. The first commercial use of Zn-coated PHS was in 2008, using the indirect process.P-20 Since then, direct forming of Zn-coated PHS has been studied. When direct formed, furnace soaking temperature and time must be controlled carefully to avoid deep microcracks.G-41, K-20 Recently developed are two new Zn-coated press hardening steel grades, 20MnB8 and 22MnSiB9-5, both reaching approximately 1500 MPa tensile strength after processing. Using grades requires a pre-cooling process after the furnace to solidify the Zn-based coating. 20MnB8 can be direct hot formed to final shape, whereas 22MnSiB9-5 can be formed in a transfer press in the “multi-step” process.K-21, H-27
Depending on the coating type and thickness, the process type, controls and investment requirements may change significantly. For example, some press hardening lines may be designed to form blanks with only Al-based coatings. Table 1 summarizes the advantages and disadvantages of several coating systems.
Table 1: Summary of coatings available for press hardening steels.
Uncoated Blanks
The earliest press hardening steels did not have any coating on them. These steels are still available and may be preferred for dry areas in automotive applications. If the steel is uncoated and the furnace atmosphere is not controlled, scale formation is unavoidable. Scale is the term for iron oxides which form due to high temperature oxidation. Scale thickness increases as the time in furnace gets longer, as seen in Figure 3. Scale has to be removed before welding, requiring a shotblasting stage. Thicker scale is more difficult and more costly to remove.M-53 Early attempts to reduce (if not avoid) scale formation saw the use of an inert-gas atmosphere inside the furnace.A-52 Today, a mixture of nitrogen (N2) and natural gas (CH4) is typically used.F-35 In China, at least one tier supplier is using a vacuum furnace to prevent scale formation.A-68
![Figure 3: Oxide layer (scale) on press hardened steel after: (a) fast resistance heating (10 seconds in air), (b) furnace heating (120 seconds in air) [REFERENCE 14]](https://ahssinsights.org/wp-content/uploads/2021/02/Fig3_Oxide-layer-on-press-hardened-steel-after-fast-resistance-heating.jpg)
Figure 3: Oxide layer (scale) on press hardened steel after: (a) fast resistance heating (10 seconds in air), (b) furnace heating (120 seconds in air).M-53
While heating uncoated steel in the furnace, if the conditions are favorable for iron (Fe) oxidation, carbon (C) may also be oxidized. When the carbon is oxidized, layers close to the surface lose their carbon content as gaseous carbon monoxide (CO) and/or carbon dioxide (CO2) is produced.S-87 The depth of the “decarburization layer” increases with dwell time in the furnace, until an oxide layer (scale) formed. Scale acts as a barrier between the bare steel and atmosphere. As the carbon is depleted in the “decarburization layer”, the hardness of the layer is decreased, as seen in Figure 4. Decarburization is usually undesirable since it lowers the strength/hardness and may negatively affect fatigue life.C-26
Figure 4: Hardness distribution of an uncoated steel after 6 minutes in a 900 °C furnace, showing hardness decrease as the surface layers lose their carbon. Image recreated after C-26.
Several methods are available to improve the corrosion resistance of uncoated PHS parts:
- E-coating after welding, before painting is a typical step of car body manufacturing, for rustproofing.
- If descaling can be done by using chromium shots (in shotblasting), a thin film of chromium-iron may grow on the surface and improve the corrosion resistance.F-14
- Vapor galvanizing (also known as Sherardizing) of uncoated steel after descaling, an experimental study described in Citation G-42.
- Electro-galvanizing after hot stamping, as described in Citation A-68.
- Change the base metal chemistry to one that is more oxidation resistant.L-60 Figure 5 compares the shiny non-oxidized surface appearance of parts made from this grade with that made from a conventional uncoated press hardening grade on the same production line with the same processing conditions.W-28
Figure 5: Oxidation resistant PHS grades may not need descaling or coatings for sufficient corrosion resistance.W-28
Aluminium-Based Coatings
The first commercially available coating on press hardening steels was patented by Sollac (now part of ArcelorMittal) in 1998. This coating was designed to address the scaling problem, but also offers some barrier corrosion resistance.C-24 The nominal coating composition is 9-10 wt.% Si, 2-4 wt.% Fe, with the balance Al.L-39 The coating may be referred to as AlSi, Al-Si, AluSi or more commonly AS. Nominal as-delivered coating thickness is 25 μm (approximately 75 g/m2) on each side, and is usually referred to as AS150, with 150 referencing the total coating weight combining both sides, expressed as g/m2. More recently, a thinner coating of 13 μm (30-40 g/m2 on each side, AS60 or AS80) is now commercially available.A-51 When AS coated blanks are “tailor rolled,” the coating thickness is also reduced in a similar percentage of the base metal thickness reduction. Corrosion protection is similarly reduced, and furnace parameters need to be adjusted accordingly.
As delivered, AS150 has a coating thickness of 20-33 μm and a hardness of approximately 60 HV. The “interdiffusion layer” (abbreviated as IDL) has a high hardness and low toughness at delivery, as seen in Figure 6a. Due to the brittle nature of the IDL, AS coated blanks cannot be cold formed unless very special precautions are taken. During heating, iron from the base metal diffuses to the coating forming very hard AlSiFe (or AlFe) layers close to surface. At the same time, Al and Si of the coating diffuse to the IDL, growing it in thickness and reducing its hardness, Figure 6b. Earlier studies have shown that heating time (and also furnace temperature) has direct effect on the final thickness of IDL, as shown in Figure 7. Once the IDL thickness surpasses approximately 16 to 17 μm, the welding current range (ΔI = Iexpulsion – Imin) may be well below 2 kA.V-15, V-21, W-34 The dwell time must be long enough to ensure proper surface roughness (see Figure 6b) for e-coatability.M-27, T-40 Figure 10 summarizes the heating process window of AS coatings. The process window may change with base metal and coating thicknesses.
Figure 6: AS coating micrographs: (a) as-delivered, (b) after hot stamping process (re-created after V-15, V-21, W-34, G-32)
![Figure 6: IDL thickness variation with furnace dwell time (Image created by REFERENCE 43 using raw data from REFERENCE 22, REFERENCE 26, and REFERENCE 27]](https://ahssinsights.org/wp-content/uploads/2021/02/Fig6_IDL-thickness-variation-with-furnace-dwell-time.jpg)
Figure 7: IDL thickness variation with furnace dwell time (Image created by B-55 using raw data from V-21, G-32, K-41.)
Hydrogen induced cracking (HIC, also known as hydrogen embrittlement) has been a major problem for steels over 1500 MPa tensile strength. AS coated steels may have higher diffusible hydrogen, when delivered, due to the aluminizing process occurring at 680 °C. In addition, AS coated grades may have a hydrogen absorption rate up to three times higher during heating.C-27 To reduce the hydrogen diffusion, it is essential to control the heating process (both heating rate and dew point in the furnace). AS coated blanks absorb hydrogen at room temperature; however, this happens at much lower rates than uncoated or Zn-coated blanks.J-21 Diffusible hydrogen can be removed from the press hardened part by re-heating the part to around 200 °C for 20 minutes or longer, in a process called de-embrittlement.V-21, G-32, G-43, J-21
For the abovementioned reasons, AS coated higher strength grades (i.e., PHS1800 and over) are required to have precise “dew point regulations” during the heating in furnace. Their final properties, especially elongation and bending angle, may be guaranteed only after bake hardening, as shown in Figure 8.B-32 Paint baking is standardized in Europe as a treatment for 20 minutes at 170 °C, which may act like a de-embrittlement treatment.E-10 Some OEMs also require dew point control and “subsequent de-embrittlement treatment” for AS coated PHS1500.
Figure 8: Effect of diffusible hydrogen (Hdiff) on mechanical properties of: (a) uncoated PHS2000, (b) AS coated PHS2000 in an uncontrolled furnace atmosphere (B-55 using raw data from C-27).
Another method to reduce the risk of hydrogen embrittlement is to adjust the coating composition. The bath chemistry for a standard AlSi coating consists of up to 90% aluminum, about 8% to 11% silicon and a maximum of 4% iron. Adding a maximum of 0.5% alkaline earth metals, like magnesium, for example has been shown to result in 40% less hydrogen diffusion into steel.R-29, T-45
Although not common in the industry, Al-Zn and Zn-Al-Mg based coatings have also been developed for press hardening processes.F-14 Recently introduced is an aluminium-silicon coating with magnesium additions. When oxidized with water vapor, Mg releases less H2 and thus may reduce the diffusible hydrogen.S-88
AS coatings may cause costly maintenance issues in roller hearth furnaces, as the coating may contaminate the rollers.B-14 Special care has to be taken to avoid the issue or prolong the maintenance intervals.
Zinc-Based Coatings
AS coatings provide some corrosion protection, known as “barrier protection”, as the coating forms a barrier between the oxidizing environment and the bare steel. It is quite common in Europe for a car to have 12 years corrosion protection warranty. To achieve such corrosion resistance, a typical car may have over 85% of its components galvanized.S-89
The use of Zn-coated PHS has been relatively low, compared to AS coated and uncoated grades. In 2015, 76% of the PHS sold in EU27+Turkey was AlSi coated. In these markets, 18% of the PHS sold was uncoated and only 6% was Zn coated.D-20 This can be attributed to the susceptibility of Zn-coated PHS to Liquid Metal Embrittlement (LME, also known as Liquid Metal Assisted Cracks (LMAC) and Liquid Metal Induced Embrittlement (LMIE)).C-28, L-46
After heating and soaking in the furnace, the base metal should be in the austenitic phase. During heating, the Zn coating reacts with the base metal and forms a thin solid layer of body-centered-cubic solid solution of Zn in α-Fe, shown as α-Fe(Zn) in Figure 9. During deformation, a microcrack can be initiated in this layer at the grain boundaries of the austenite in the base metal, as indicated in Figure 9a. As the crack propagates, zinc from the α-Fe(Zn) layer diffuses along the austenite grain boundary and combines with iron from the base steel to form additional α-Fe(Zn), Figure 9b. Cracks propagate through the weak a-Fe(Zn) grain boundary layer, allowing liquid zinc (with diffused iron) to advance into the capillary crack (Figure 9c). After quenching, the base metal transforms to martensite and the liquid Zn transforms to a hard and brittle intermetallic phase, Γ-Fe3Zn10.C-28
Figure 9: Schematic illustration of microcrack formation. (re-created based on C-28.)
To avoid LME, three methods can be employedK-20:
- Forming in the absence of liquid Zn,
- Reducing stress level,
- Reducing material susceptibility.
There are no breakthroughs to address the last two items. Forming a part in the absence of liquid Zn involves either of two process routes: (1) Indirect press hardening (also known as form hardening), or (2) Pre-cooled direct processes.
In the direct forming of Zn-coated blanks, with or without pre-cooling, microcracks in the base metal may be observed. Microcracks less than 10 μm into the base metal does not affect the fatigue strength of the part.K-20 Microcrack depth is a function of coating thickness, furnace conditions (temperature and dwell time, see Figure 10), forming severity and forming temperature. It may be possible to direct form galvannealed (GA coated) blanks.
The boiling point of pure zinc (907 °C) is very close to the austenitization temperature of 22MnB5 (885 °C), so the heating process window of Zn-coated blanks must be controlled precisely. When the furnace dwell time is too short, deeper microcracks may be observed. When the furnace dwell time is too long, corrosion performance may be degraded. Thus, heating process window of Zn-coated blanks is significantly narrower than that of AS-coated blanks.B-14, S-90
Figure 10: Heating process window of AS and Zn coatings (representative data, may not be accurate for all sheet and coating thicknesses, re-created based on B-14, S-90.
Zn-based coatings may result in very low diffusible hydrogen after press hardening. In one studyJ-21, no diffusible hydrogen was detected, as long as the furnace dwell times are shorter than 6 minutes. Even after 50 minutes in the furnace, diffusible hydrogen was found to be around 0.06 ppm. Zn coatings do not act as a barrier for hydrogen desorption (losing H through the surface). Even at room temperature, Zn coated blanks may lose most of the diffusible hydrogen within a few days (also referred to as aging).
![Figure 10: Evolution of galvanized coating: (a) as delivered: Ferrite+Pearlite in base metal, almost pure Zn coating with Al-rich inhibition layer, (b) at high temperatures: austenite in base metal + α-Fe(Zn) and liquid Zn + surface oxides, (c) after press hardening: martensite in base metal + α-Fe(Zn) and Γ-phase coatings + surface oxides. The oxides are removed prior to welding and painting [REFERENCE 30]](https://ahssinsights.org/wp-content/uploads/2021/02/Fig10_Evolution-of-galvanized-coating.jpg)
Figure 11: Evolution of galvanized coating: (a) as delivered: Ferrite+Pearlite in base metal, almost pure Zn coating with Al-rich inhibition layer, (b) at high temperatures: austenite in base metal + α-Fe(Zn) and liquid Zn + surface oxides, (c) after press hardening: martensite in base metal + α-Fe(Zn) and Γ-phase coatings + surface oxides. The oxides are removed prior to welding and painting.J-21
Zn-based coatings may have a yellowish color after hot stamping. The surface oxides have to be removed before welding. This is typically done by shotblasting.
PHS blanks with a ZnNi coating were previously available. The ZnNi coating provided a low friction coefficient, a large process window in the furnace, the ability to be cold formed (indirect or two-step hybrid processes were also possible) and decreased susceptibility to LME.B-56 ZnNi coated PHS was used in the rear rail of the Opel Adam city carH-57 for a short period, until the coating was discontinued.C-29
Varnish Coatings
Another method to avoid scaling and decarburization is to apply varnish coatings. In this method, uncoated steel can be either coil coated or blanks can be manually coated with the paint-like varnish coatings.B-14 These coatings may also be known as “paint-type” or “sol-gel”.
![Figure 11: Manual application of a varnish coating. [REFERENCE 7]](https://ahssinsights.org/wp-content/uploads/2021/02/Fig11_Manual-application-of-a-varnish-coating.jpg)
Figure 12: Manual application of a varnish coating.F-34
Depending on the type of coating, they may allow very fast heating – including induction and conduction heating with electric current. Since the coating does not require time to diffuse, furnace heating may be completed in less than 2 minutes.F-34 Again, depending on the type, surface conditioning may not be required before welding or e-coating.B-14
They were used in automotive industry between 2005 and 2010. By 2015 there were four different types of varnish coatings, some of which are now discontinued.B-14 These coatings may be useful for prototyping and low volume production.
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Citations
Citation:
Z-15. D.J. Zhou, and K. Kannan, “The Effect of Combination Beads on Springback: Experimental Study & Virtual Study“, Presented at 2021 Great Designs in Steel, Sponsored by American Iron and Steel Institute.
Press Requirements
AHSS products have significantly different forming characteristics and these challenge conventional mechanical and hydraulic presses. The dramatically higher strength of these new steels result in higher forming loads and increased springback. Higher contact pressures cause higher temperatures at the die-steel interface, requiring high performance lubricants and tool steel inserts with advanced coatings. Although the decrease in ductility is not as severe as seen with HSLA grades of a similar strength level, there is a reduction in formability. Further complicating matters is that in addition to traditional stamping failures due to necking when the part strains exceed the forming limit curve, AHSS grades have additional failure modes such as reduced bendability and cut edge ductility (collectively called local formability failure modes) which are not predicted using current analytical techniques.
These challenges lead to issues with the precision of part formation and stamping line productivity. The stamping industry is developing more advanced die designs as well as advanced manufacturing techniques to help reduce fractures and scrap associated with AHSS stamped on traditional presses. Using a servo-driven press is one approach to address the challenges of forming and cutting AHSS grades. Recent growth in the use of servo presses in the automotive manufacturing industry parallels the increased use of AHSS in the body structure of new automobiles.
Characteristics of Servo Presses
A servo press uses a servomotor as the drive source. Servo press systems are more flexible than flywheel-driven presses and are both faster and more accurate than hydraulic presses. A servomotor allows for control of the position, direction, and speed of the output shaft in contrast to a constant cycle speed of flywheel driven presses, for example. New forming techniques take advantage of this flexibility, achieving more complex part geometries while maintaining dimensional precision.
Mechanical presses are powered by an electric motor that drives a large flywheel. The flywheel stores kinetic energy, which is released through various drive types like cranks, knuckle joints, and linkages. Powering hydraulic presses are electric motors which drive hydraulic cylinders to move the ram up or down.
Servo-driven presses can be either mechanical or hydraulic. In servo-mechanical presses, the high-powered servo motor allows for direct driving of the mechanical press without using a flywheel and clutch. Up to the rated speed, maximum torque is available. Beyond this rated speed, the available torque decreases until reaching the maximum speed. If forming speeds remain below this rated speed, servo presses may have an advantage over flywheel driven presses since full tonnage is available even at lower strokes per minute. This is a useful feature if heat build-up limits how fast the part is capable of running.
Traditional hydraulic presses use variable volume pumps powered by constant velocity electric motors. Servo-hydraulic presses either combine conventional electric motors with servo (proportional) valves or pair servo motors with simple pumps and valves. Typically, servo-hydraulic presses reach higher slide speeds than conventional hydraulic presses, but usually are not faster than servo-mechanical presses over a complete cycle. Due to this advantage, most automotive stampers use servo-mechanical presses.
For both servo-mechanical and servo-hydraulic presses, the other press components remain the same as conventional presses. Figure 1 compares pertinent elements of a flywheel-driven mechanical press with one driven by a servo motor.
Figure 1: Drive components in a mechanical press. A) Flywheel driven; B) Servo motor driven.A-9
Advantages of Servo Presses
Servo press technology has many advantages compared to flywheel-driven mechanical presses when working with AHSS materials. Press manufacturers and users claim advantages in stroke, speed, energy usage, quality, tool life and uptime; these of course are dependent upon part shape and forming complexity. Figure 2 shows the difference between the available motions of flywheel-driven mechanical presses verses servo driven mechanical presses. The slide motion of the servo press can be programmed for more parts per minute, decreased drawing speed to reduce quality errors, or dwelling or re-striking at bottom dead center to reduce springback.
Figure 2: Comparison of Press Signatures in Fixed Motion Mechanical Presses and Free-Motion Servo-Driven Presses.M-3
Some examples highlighting the flexibility of servo press motions include:
- Adjustable Stroke Length: The servo press has an adjustable stroke; the programmable slide motion allows for customizing the stroke profile to suit the characteristics of the part. By changing the programmed profile, the same press can produce deep draw stampings, shallow part stampings, or even blanking. This flexibility allows metal formers to use servo presses for multiple job types.
- Adjustable Stroke Direction – Pendulum Motion: Unlike in flywheel-driven presses, servo-driven presses do not need to return to Top Dead Center. A servomotor can change the rotation direction at predefined angles (like restricting the rotation per stroke between 90 degrees and 270 degrees) to increase production rates. This is known as pendulum mode.
- Adjustable Stroke Direction – Attach-Detach (forward/reverse) Motion: The slide moves downward at a pre-set speed, reaches a predetermined depth, and rises slightly. The slide then moves downward again, and the system repeats this motion. When the slide motion changes from downward to upward, the forming forces become negligible and the elastic deformation of dies and machine recover. This has the potential to mitigate springback that is associated with the forming of higher strength materials in traditional mechanical presses. By holding pressure at the bottom of the stroke, releasing the pressure, and then reapplying it, multiple stamping hits can be made in just one cycle, setting the form of the product, and eliminating the need for secondary operations.
- Improved Cycle Time: Servo-driven presses achieve improved cycle times compared with standard mechanical presses, improving stamping plant productivity. This comes from the ability of servo presses to run faster on the approach and retract portions of the cycle when no work is performed on the sheet metal. Figure 3 shows cycle rates for comparable stroke heights for both a servo-mechanical press and a traditional flywheel-driven mechanical press in which the stroke-time signature takes the form of a sine wave. This example shows a 60% improvement in parts per minute. Quality likely improved also, for reasons discussed below.
Figure 3: Cycle rates for servo-driven and flywheel-driven mechanical presses.B-9
- Variable Speed Slide and Acceleration/Deceleration: At any part of the press stroke, the slide speed in the forming stage can be programmed to accelerate, decelerate, or remain constant, as necessary. Decreasing the touch speed of upper and lower dies minimizes impact shocks, improving quality while increasing the life of tools and presses. After forming, returning to the top slide position (not necessarily Top Dead Center) can occur quickly. These slide movements may alter the frictional characteristics between dies and sheet materials.
- Dwell Function of the Slide: Programming the slide to dwell at bottom dead center (BDC) is critical when stamping press hardenable steels, as this time allows for the necessary cooling under full press load that is key to achieving the dimensional and strength properties. In other applications, coining pressure at BDC ensures shape and dimensional accuracy of the product formed.
- Quality Improvements: High cycle rates in flywheel driven presses risk heating the sheet metal and tooling, leading to breakdown of lubricants and accelerated tool wear. During blanking or hole piercing, high cycle rates increase the reverse-tonnage snap through loads, damaging cut edge quality and sending shock waves through your tooling and press. Servo-driven presses can slow down just before contact with the sheet metal to dramatically reduce impact forces and shock loads, yet maintain high productivity by having the ram move much faster while no work is occurring. Improved product consistency should translate to fewer rejections.
- Longer Tool Life: Decreasing the tool impact speed reduces impact loading, thus maximizing tool life. With servo-driven presses, there is not a productivity penalty. Slower speeds and less aggressive forming conditions may allow for use of a lower grade lubricant. Pulsating or oscillating slide motion can extend the working limit.
- Uptime: Fewer moving parts and reduced shock loads translate into reduced die fatigue and greater uptime.
- Independent Motion of the Die Cushion: Programmable die cushions available with some servo presses function like hydraulic cushions found on some mechanical presses. A pressure sensor within the servo die cushion controls its position. This allows for the optimization of metal flow in the flange between the die and the blank holder. Varying the blank holder force during the stroke reduces springback in AHSS.S-44 Additionally, servo-driven die cushions can regenerate energy when the upper die and slide push the cushion downward.
- Formability Improvements: Using a servo press with a modified press cycle has allowed some stampers to use a less expensive, less formable grade without increasing rejects or compromising dimensional tolerance.
- Reduced Noise: Fewer moving parts combined with slower tool impact speeds and reduced snap-through loads lead to a quieter work environment.
- Synchronization with Transfer and Feed Systems: Servo presses provide for adjustments in slide position and speed, allowing for optimized timing to coordinate with part transfer systems.
- Multiple Operations in One Cycle: Control of the slide position combine with the ability to dwell make In-die operations possible.
- Energy Savings: During operations like blanking or drawing, servomotor power is used only while the press is moving, unlike in mechanical presses with a continuously rotating flywheel and clutch/brake mechanism. These mechanical press components each have energy losses due to friction. The lower energy consumption and associated costs seen in servo-driven presses are even more substantial in larger capacity presses. Also contributing to the energy savings is the dynamic braking operation of the servo driven motor, where the energy from braking transfers back into the power system. Similarly, compressed die cushions can feed energy to the storage device. When economically justified, installation of an external energy storage unit can make up for energy peaks while reducing the nominal power drawn from the local power supply system.
Figure 4 illustrates power storage and output in a servo-mechanical press system over the course of a cycle. In this example, the press operates with two main motors, each having a maximum output of 175 kW. An external energy device stores energy from the slide deceleration, and is tapped when the press motion requires more than 175 kW from each motor. The stored energy (maximum of 350 kW combined in the two 175 kW motors) is available during peak power requirements, enabling the facility power load to remain nearly constant at around 50 kW.B-9
Figure 4: Power storage and output in a servo mechanical press system.B-9
Press Force and Press Energy
Categorizing both mechanical and hydraulic presses requires three different capacities or ratings – force, energy, and power. Historically, when making parts out of mild steel or even some HSLA steels, using the old rules-of-thumb to estimate forming loads was sufficient. Once the tonnage requirements and some processing requirements were known, stamping could occur in whichever press met those minimum tonnage and bed size requirements.
In these cases, press capacity (for example, 1000 kN) is a suitable number for the mechanical characteristics of a stamping press. Capacity, or tonnage rating, indicates the maximum force that the press can apply without damaging its components, like the machine frame, slide-adjusting mechanisms, pitman (connection rods) or main gear bushings.
A servo press transmits force (not energy) the same way as the equivalent conventional press (mechanical or hydraulic). However, the amount of force available throughout the stroke depends on whether the press is hydraulic or mechanically driven. Hydraulic presses can exert maximum force during the entire stroke as tonnage generation occurs via hydraulic fluid, pumps, and cylinders. Mechanical presses exert their maximum force at a specific distance above bottom dead center (BDC), usually defined at 0.5 inch. At increased distances above bottom dead center, the loss of mechanical advantage reduces the tonnage available for the press to apply. This phenomenon is known as de-rated tonnage, and it applies to conventional mechanical presses as well as servo-mechanical presses. Figure 5 shows a typical press-force curve for a 600-ton mechanical press. In this example, when the press is approximately 3 inches off BDC, the maximum tonnage available is only 250 tons – significantly less than the 600-ton rating.
Figure 5: The Press-Force curve shows the maximum tonnage a mechanical press can apply based on the position of the slide relative to the bottom dead center reference distance. This de-rated tonnage applies to both conventional mechanical presses as well as servo-mechanical presses.E-2
Press energy reflects the ability to provide that force over a specified distance (draw depth) at a given cycle rate. Figure 6 shows a typical press-energy curve for the same 600-ton mechanical press. The energy available depends on the size and speed of the flywheel, as well as the size of the main drive motor. As the flywheel rotates faster, the amount of stored energy increases, reflected in the first portion of the curve. The cutting or forming process consumes energy, which the drive motor must replenish during the nonworking part of the stroke. At faster speeds, the motor has less time to restore the energy. If the energy cannot be restored in time, the press stalls. The graph illustrates how the available energy of the press diminishes to 25% of the rated capacity when accompanied by a speed reduction from 24 strokes per minute to 12 strokes per minute.
Figure 6: A representative press-energy curve for a 600 ton mechanical press. Reducing the stroke rate from 24 to 12 reduces available energy by 75%.E-2
Rules of thumb are useful to estimate press loads. However, a better evaluation of draw force, embossment force, and blank holding force comes from simulation tools. Many programs enable the user to specify all of the system inputs. This is especially important when forming AHSS because the rapid work hardening seen in these grades has a major effect on the press loads. In addition, instead of using a simple restraining force on blank movement, incorporating the geometry and effects of the actual draw beads leads to improved simulation accuracy.
A common shortcut taken in simulation is the assumption that the tools are rigid during forming. In practice, however, tools will deform elastically. Further increasing this deflection of the dies (sometimes called breathing) is the higher work hardening of AHSS grades. This discrepancy leads to a significant increase in the determined press loads, especially when the punch is at home position. Hence, for a given part, the draw depth used for the determination of the calculated press load is an important parameter. Applying the nominal draw depth may result in an over-estimation of press loads. Similarly, assuming that the structure, platens, bolsters, and other components of the press are completely rigid may lead to variation in press loads, especially after moving the physical tooling from one press to another.
In all cases, validation of all simulation predictions is good practice. Every simulation contains assumptions, with some being more critical than others. Use practical stamping tests to determine the optimum parameter settings for the simulation. Quick items to confirm simulation matches with reality include draw-in amounts (lay draw panel on top of blank) and press tonnages (check load monitors). When physical panels match simulation results, confidence in the simulation accuracy rises.
In processes like restrike operations, simulation may not accurately estimate press loads. In these cases, run “what-if” simulations to observe forming trends for a given part, which helps to develop a more favorable forming-process design.
Setting Draw Beads with Air Cushions and
Nitrogen Cylinders
Air cushions and nitrogen cylinders are common on single action mechanical presses to give them a double action. There is a tonnage spike on initial contact with the blankholder and setting of the draw beads, occurring while the press is still several inches off Bottom Dead Center. This spike increases the likelihood of a mechanical press being damaged. Figure 7 shows the potential negative consequences when a mechanical press exceeds the rated capacity of the press and the associated components.
Figure 7: Broken connecting rod on a mechanical press.M-5
A nitrogen die cushion in a single-acting press needs to apply considerable force to set draw beads in AHSS sheets before drawing begins, as well as to apply stake beads at the end of the stroke for springback control. In some cases, binder separation may occur because of insufficient cushion tonnage, resulting in a loss of control for the stamping process and excessive wrinkling of the part or addendum. The high impact load on the cushion may occur several inches up from the bottom of the press stroke, where de-rated tonnage means a reduced maximum load to avoid press damage. Flywheel-driven mechanical presses are susceptible to damage due to these shock loads, since the impact point in the stroke occurs when the press is travelling at a higher velocity. The high shock loads dissipate additional flywheel energy well above bottom dead center of the stroke. Therefore, a nitrogen-die cushion may be inadequate for optimum pressure and process control when working with AHSS.
Staggering the heights of the nitrogen cylinders so they do not all engage at the same time is one way to reduce the shock load (Figure 8). A double-action press will set the draw beads when the outer slide approaches bottom dead center where the full tonnage rating is available and where the slide velocity is substantially lower. This minimizes any shock loads on the die and press with resultant load spikes less likely to exceed the rated press capacity.
Figure 8: Staggered nitrogen cylinders reduce the initial shock load when setting draw beads by engaging at different depths in the press stroke.M-5
Press Considerations When Working With Higher Strength Steels (U-13)
The increased forces needed to form, cut and trim higher-strength steels create significant challenges for pressroom equipment and tooling. These include excessive tooling deflections, damaging tipping-moments, and amplified vibrations and snapthrough forces that can shock and break dies—and sometimes presses. Stamping AHSS materials can affect the size, strength, power and overall configuration of every major piece of the press line, including material-handling equipment, coil straighteners, feed systems and presses.
Here is what every stamper should know about higher-strength materials:
- Because higher-strength steels require more stress to deform, additional servo motor power and torque capability may be needed to pull the coil material through the straightener. Additional back tension between the coil feed and straightening equipment also may be required due to the higher yield strength of the material in the loop as the material tries to push back against the straightener and feed system.
- Because higher-strength materials require greater stress to blank and punch as compared to HSLA or mild steel, they generate proportionally increased snapthrough and reverse-unloading forces. High-tensile snapthrough forces introduce large downward accelerations to the upper die half. These forces work to separate the upper die from the bottom of the ram on every stroke. Insufficient die-clamping force could cause the upper-die half to separate from the bottom of the ram on each stroke, causing fatigue to the upper-die mounting fasteners.
- Because energy is expended with each stroke of the press—and this energy must be replaced—critical attention must focus on the size (horsepower) of the main drive motor and the rotational speed of the flywheel in higher-strength-steel applications. The main motor, with its electrical connection, provides the only source of energy for the press and it must generate sufficient power to meet the demands of the stamping operation. The motor must be properly sized to replace the increased energy expended during each press stroke. For these reasons, some stampers consider the benefits of servo-driven presses for these applications.
As steels becomes stronger, a corresponding increase in process knowledge is required in terms of die design, construction and maintenance, and equipment selection.
Case Study: The Importance of Press Energy
in Flywheel Driven Presses
In a flywheel-driven mechanical press, the size of the main motor, flywheel mass, and rotation speed of the flywheel become critical. The main motor, along with its electrical connections, is the only source of energy for the press and it must have sufficient power to supply the demands of the stamping operation. As the flywheel is an energy storage device, it must be able to store and deliver the required energy when needed. The stored energy varies by the square of the speed; thus, flywheels can store a large amount of energy when the press is running at full speed. If heat generation and forming problems occur when stamping AHSS grades, operators may be inclined to slow down press speeds. However, this slowdown may lead to not meeting the energy requirements to form the part, ultimately resulting in the press stalling.
Take the example of a part with a draw depth of 2 inches. If stamped from HSLA 350/450, it may need 150 tons of forming force, for a total of 300 inch-tons of forming energy (2 inches times 150 tons). On this press, 14 strokes per minute is sufficient to generate enough forming energy, as indicated by the green dot in Figure 9.
Studies have shown that the forming tonnage of DP steels may be twice that of HSLA steels, which means that the 2 inch draw depth could need 300 tons of forming force, for a total of 600 inch-tons of forming energy. The red dot in Figure 9 shows that our press must run at 20 strokes per minute to ensure that there is sufficient forming energy. The speed is well within the capability of the press, but our strategy of running faster may not be appropriate for forming AHSS grades where heat generation could result in lubricant breakdown leading to die wear, galling, and scoring. Running the press slower to avoid these concerns increases the risk of stalling.
Figure 9: Higher strength materials requiring increased forming energy also require faster cycle times. If the greater cycle time cannot be maintained due to reasons like heat buildup or lubricant breakdown, a flywheel-driven press may stall.E-2
Case Study: Press Force and Press EnergyH-3
Predicting the press forces needed initially to form a part is known from a basic understanding of sheet metal forming. Different methods are available to calculate drawing force, ram force, slide force, or blankholder force. The press load signature is an output from most forming-process development simulation programs, as well as special press load monitors.
Most structural components include design features to improve local stiffness. Typically, forming of the features requiring embossing processes occurs near the end of the stroke near Bottom Dead Center. Predicting forces needed for such a process is usually based on press shop experiences applicable to conventional steel grades. To generate comparable numbers for AHSS grades, forming process simulation is recommended.
In Citation H-3, stamping simulations evaluated the forming of a cross member having a hat-profile with an embossment formed at the end of the stroke (Figure 10). The study simulated press forces and press energy involved for drawing and embossing a channel section from four steel grades approximately 1.5 mm thick: mild steel, HSLA 250/350, HSLA 350/450, and DP 350/600. Figures 11A and 11B clearly show that the embossing phase rather than the drawing phase dominates the total force and energy requirements, even though the punch travel for embossing is only a fraction of the drawing depth.
Figure 10: Cross section of a component having a longitudinal embossment to improve local stiffness.H-3
Figure 11: Embossing requires significantly more (A) force and (B) energy than drawing, even though the punch travel in the embossing stage is much smaller.H-3
Figures 12 and 13 highlight the press energy requirements, showing the greater energy required for higher strength steels. The embossment starts to form at a punch displacement of 85 mm, indicated by the three dots in Figure 12. The last increment of punch travel to 98 mm requires significantly higher energy, as shown in Figure 13. Note that compared with mild steel, the dual phase steel grade requires significantly more energy to form the part to home with the 98 mm travel.
Figure 12: Energy needed to form the component increases for higher strength steel grades. Forming the embossment begins at 85 mm of punch travel, indicated by the 3 dots.H-3
Figure 13: Additional energy required to form the embossment increases for higher strength steel grades.H-3
It is not only embossments that require substantially more force and energy at the end of the stroke. Stake beads for springback control engage late in the stroke to provide sidewall stretch. Depending on the design of the forming process, the steel into which the stake beads engage may have passed through conventional draw beads for metal flow control, and therefore are work-hardened to an even higher strength. This leads to greater requirements for die closing force and energy. Certain draw-bead geometries which demand different closing conditions around the periphery of the stamping also may influence closing force and energy requirements.
Energy Requirements
Using existing production data can estimate press loads for simple geometries, knowing just the thickness and tensile strength. There is a proportional increase in forming load (F) resulting from the product of thickness (t) and tensile strength (Rm). For example, in your current production (1), you know your drawing force F1, and the thickness t1 and tensile strength Rm1 of the steel you are forming. You are switching production (2) to a new tensile strength Rm2 and thickness t2, and you want to know the drawing force F2. Set up the following equation:
which is the same as
The work in Citation T-11 studied the press loads required to form a cross member with a simple hat profile from two steels of the same 0.7 mm thickness: HSLA 350/450 and DP 300/500 steels. Here, it was known that an HSLA coil with 433 MPa tensile strength required 791 kN of drawing force. Of interest was the drawing force required to stamp a dual phase steel with a tensile strength of 522 MPa. Using the equations above, the estimated force F2 was calculated as 954MPa:
This estimated 954 MPa compares favorably with the force measured during their trial stamping of 934 MPa. Using this technique typically is sufficient to estimate the press requirements, but may lead to an over-estimation of the actual loads.
The energy required for plastically deforming a material (force times distance) has the same units as the area under the true stress-true strain curve. For this reason, assessing the forming energy requirements of two grades requires comparing the respective areas under their true stress – true strain curves. The shape and magnitude of these curves are a function of the yield strength and work hardening behavior as characterized by the n-value. At the same yield strength, a grade with higher n-value will require greater press energy capability, as highlighted in Figure 14 which compares HSLA 350/450 and DP 350/600. For these specific tensile test results, there is approximately 30% greater area under the DP curve compared with the HSLA curve, suggesting that forming the DP grade requires 30% more energy than required to form a part from the HSLA grade.
Figure 14: True stress-strain curves for two materials with equal yield strength.T-11
Greater work hardening of DP steels results in higher forming tonnage requirements when compared to HSLA grades at the same incoming yield strength and sheet thickness. However, AHSS applications justify a thickness reduction, and along with this is a reduction in the required press load. The required power is a function of applied forces, the displacement of the moving parts, and the speed. The energy rating of a press is also a function of applied press load and the distance over which the load is applied. For example, pushing 200 tons through 3 inches of deep drawing requires 600 inch–tons of energy. Changing the part to AHSS could require 500 tons of force working through the same 3 inch distance, requiring 1500 inch-tons of energy. Each stroke of the press expends a given amount of energy, all of which must be replaced before the next stroke begins.
Key Points
- Forming AHSS steels require increased press loads primarily because of their increased work hardening and increased strength levels.
- Press energy availability is as important as press force capacity when selecting a press to form AHSS grades.
- Slowing a press to counteract heat build-up may lead to stalling of flywheel-driven presses due to insufficient press energy availability.
- Comparing the areas under the respective true stress-true strain curves is one way to assess the additional increase in required press energy between two grades.
- Embossments and other features engaging off bottom dead center require sufficient de-rated tonnage and press energy at appropriate distances off bottom.
- DP 350/600 requires about twice the energy to form a hat-profile cross member than the same cross member formed from mild steel.
- Forming and process simulations are the most efficient and accurate approach to estimating the punch force and energy requirements to process AHSS grades.
- Servo-driven presses have characteristics which make them suitable for forming complex geometries using advanced high strength steels. Adding to their business case, servo-driven presses are more energy efficient and the flexibility in press cycle profiles can lead to enhanced productivity.
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Testing and Characterization
The term local formability describes when part and process design, in addition to sheet metal properties like strength and elongation, influence the amount of deformation the metal can undergo prior to failure. Cutting, punching or other methods of obtaining a trimmed blank or an internal hole results in cracks, rough edges, work-hardening and other edge damage – all of which influences edge quality. The challenges of capturing all of the factors that influence edge quality makes the prediction of fracture severity and cut edge expansion very difficult and usually impossible. The many variables highlight the need for a standardized test method. However, restricted sample preparation and testing variables in these standards do not reflect the variety of conditions encountered in production environments. Use caution when comparing results generated under different conditions.
Hole Expansion Testing
The Hole Expansion test (HET) quantifies the edge stretching capability of a sheet metal grade having a specific edge condition. Higher values of the hole expansion ratio are associated with grades and forming methods more likely to have improved local formability characteristics.
Steel producers study hole expansion capacity to create new products with targeted edge stretching performance through modifications of chemistry, rolling and thermal practices. Product designers use the hole expansion test to determine if the chosen steel grade has the inherent forming characteristics to meet their targeted shape with their chosen forming system. If they are not compatible, the chosen grade must change or aspects of the forming process must change, or possibly both.
ISO 16630 is the primary standard used which describes the test method and constraints.I-9 Others, like JIS Z 2256J-6 are based on the ISO standard, with only minor differences, if any. This standard specifies use of a 10mm diameter hole created with a 12% clearance. The sample containing the hole is clamped in place, and a conical punch having a 60 degree apex angle expands that initial hole (Figure 1). The test stops after observation of a through-thickness crack or upon experiencing a load-drop exceeding a critical threshold (Figure 2). The hole expansion ratio (HER), also known as the Hole Expansion Capacity (HEC), is simply the percent expansion of the diameter of the initial hole, typically shown as the Greek letter lambda, λ.
Figure 1: Schematic of Hole Expansion Test.A-10
Figure 2: Expanded Edge at the end of a Hole Expansion Test performed using a conical punch. The arrow points to the through-thickness crack that ended the test.E-2
The sample preparation and testing requirements of ISO 16630 are well-defined for good reason. Factors known to influence the hole expansion ratio include:
Even with these rigorously defined procedures, the test results can be heavily influenced by specimen preparation technique, specific test parameters, and human subjectivity – in other words, poor gage R&R (repeatability and reproducibility). For example a group of European steel researchers reported “an unacceptably large difference between labs” with regard to hole expansion testing. They ultimately concluded that the “difference is too large for the method to be useful in practice”.
A-76
Testing sheet steels of different thicknesses in a laboratory setting requires having multiple punches and/or dies of different diameter to maintain a consistent clearance, which is based on a percentage of the sheet thickness tested.
In production, the punch-to-die clearance can change during the life of the part, both from tooling wear as well as press misalignment. There is the additional risk that clearance can vary around the perimeter of the cut section, leading to inconsistent performance. Increasing sheet metal strength magnifies this issue.
The method used to create the free edge influences the edge quality. Improved edge quality and reduced mechanical work hardening of the edge is achieved by laser cutting, EDM cutting, water jet cutting, or fine blanking processes, and will typically improve the hole expansion ratio. Trim steel clearances, shear angles, tool steel types, and sharpness also impact hole expansion test results.
In the example shown in Figure 3, the hole expansion ratio is reduced from 280% for a milled or water jet edge down to 80% for a traditional cut edge. If clearances further increase – which could happen without proper tooling maintenance over the life of the part – the ability to expand a cut edge further decreases.
Figure 3: Hole Expansion Capacity Decreased as Edge Quality Decreases. (Based on data from Citation H-1.)
Figure 4 highlights the effect of punched vs machined holes, showing the edge damage from punching lowers the hole expansion capability. This edge damage becomes a key component of what is known as the Shear Affected Zone, or SAZ. DP steels and TRIP steels have a large hardness difference between the constituent phases, and therefore are associated with lower hole expansion ratios than HSLA and CP steels, where the phases are of more similar hardness. The influence of the metallurgical phase hardness difference is explored here. Detailed studies of sheared edge stretchability as a function of clearance, edge preparation, and grade are shown in Citations K-6 and K-10.
Figure 4: Hole expansion test results comparing punched and machined holes showing effect of damage to edge stretchability. (Based on data from Citation V-1.)
Over time, the targeted edge quality degrades and targeted clearance changes without proper attention. A study documented in Citation C-1 evaluated the hole expansion ratio created by hole punching tools as they wore in a production environment. Tools evaluated were made from 60 HRC uncoated Powder Metallurgy tool steels. Data in Figure 5 show the percent hole expansion from newly ground punches and dies (Sharp Tools) and from used production punches and dies (Worn Tools). The radial clearance was 0.1 mm. A rust preventative oil was applied to the steels during the punching; a lubricant oil was applied during hole expansion. Tool wear and possible micro-chipping resulted in a poor edge condition. The clearance was not significantly affected, but the steel edges suffered cold work which dramatically affected their hole expansion results.
Figure 5: Impact of production tooling condition on hole expansion performance. (tests conducted w 50 mm diameter conical punch).C-1
Conclusions from Citation C-1 include:
- The best quality edge condition will produce the best results
- Tooling must remain sharp and damage-free to maintain the consistency in edge conditions.
- The burr should be in contact with the punch rather than on the freely-expanding side
- Hard and wear resistant tools, such as those produced from coated powder metallurgy (PM) tool steels, are highly recommended.
Additional information on tool materials can be found here and other articles in that category.
The ISO 16630 specificationI-9 eliminates one variable by prescribing the use of a 10 mm diameter hole, but it is important to understand that starting hole diameter influences the degree to which that hole can be expanded. A study that included mild steels to AHSS grades evaluated the effect of starting hole diameter.I-10 All steels were 1.2mm, punched with a clearance of 12.5%, and expanded with a conical punch having a 60° apex angle. As the starting diameter increases, the degree to which the hole can be expanded decreases, Figure 6. Note that as the strength increases, this effect appears to be minimized.
Figure 6: Hole Expansion Ratio Decreases as Initial Hole Diameter Increases.I-10
Increasing the starting hole diameter may help to distinguish between different grades.K-11 Similar hole expansion performance exists between DP980 and TRIP780 under ISO 16630 test conditions (punched 10 mm hole). It is easier to discern better performance in the TRIP780 product when performing a similar test with a 75 mm diameter punched hole (Figure 7).
Figure 7: Effect of Initial Punched Hole Diameter on Hole Expansion. (Based on Data from Citation K-11.)
The position of the burr relative to the punch affects performance in a hole expansion test. Detrimental effects of an expanding edge are minimized If the burr is on the punch side. Having the burr on the punch side, rather than the freely expanding side, minimizes the detrimental effects of the expanding edge. The primary reason is the outer surface is in a greater degree of tension than the surface next to the punch.
Figure 8 examines the effect of edge condition and clearance on DP 590 expanded with a conical punch.K-10 The data suggests that there could be up to a 20% increase in sheared edge extension capability just related to the burr position on holes punched with conventional clearances. This should be considered in die processing materials and designs sensitive to edge expansion.
Figure 8: The Effect of Burr Orientation on Hole Expansion as a Function of Clearance on DP590. “Burr Up” means away from the punch; “Burr Down” means in contact with the punch.K-10
Shown in Figure 9 is the influence of burr orientation and material grade.K-10 The 50XK grade shown is HSLA 350Y/450T, where there is a significant improvement in the measured hole expansion related to the position of the burr relative to the punch. The magnitude of this difference decreases as strength increases, but persists for all grades tested.
Figure 9: The Effect of Burr Orientation on Hole Expansion as a Function of Different High Strength Steel Grades “Burr Up” means away from the punch; “Burr Down” means in contact with the punch.K-10
The shape of the punch used to expand the hole impacts the degree to which it can be expanded. Figure 10 shows generalizations of the three most-common shapes: a conical punch, a flat punch, and a hemispherical punch.
Figure 10: Sketches of Punches Used for Hole Expansion: Conical, Flat, and Hemispherical.
Metal motion and appearance changes depending on the type of punch used. Using a conical punch leads to the shape shown in Figure 11a, with a flat punch leading to the appearance shown in Figure 11b.S-3 The operations are sometimes described as hole expansion when accomplished with a conical punch, and hole extrusion with use of a flat punch.
Figure 11a: Sample appearance after testing with conical punch.S-3
Figure 11b: Sample appearance after testing with flat punch.S-3
The ISO 16630 hole expansion test specifies the use of a conical punch with a 60 degree apex angle. Here, the free edge undergoes stretching and bending. Using a flat punch instead of a conical punch eliminates the bending component, and all deformation is from only edge stretching. These strain state differences lead to different sheared edge extension performance, with greater expansion prior to cracking achieved with holes expanded using a conical punch. This improved performance with conical rather than flat punches has been attributed to the presence of the bending component.N-10 Edge condition does not appear to influence hole expansion capability when a flat bottom punch is used.
Shown in Figures 12 to Figure 15 are the effects of burr orientation and punch type, which vary as a function of metal grade. Figure 16 compares the performance of reamed holes when expanded with either conical or flat punches. Where the tested grades perform similarly when expanded with a flat punch, the conical punch leads to exceptional performance of reamed holes of 3 of the 4 grades. The relatively poor performance of the DP780 grade may be due to the hardness differences between the ferrite and martensite components, noting that there is more martensite in DP780 than DP 600. In the study from which the data was taken, the complex phase steels had a yield/tensile ratio of approximately 87%, while for the dual phase grades the yield/tensile ratio was approximately 60%.P-13
Figure 12: Effect of Burr Orientation on Hole Expansion from a Conical Punch. (Based on Data from Citation P-13.)
![Figure 13: Effect of Burr Orientation on Hole Expansion from a Flat Punch [Based on Data from Reference 11]](https://ahssinsights.org/wp-content/uploads/2020/07/Pasted-into-Hole-Expansion-Testing-Intro-31.jpg)
Figure 13: Effect of Burr Orientation on Hole Expansion from a Flat Punch. (Based on Data from Citation P-13.)
![Figure 14: Effect of Punch Type on Hole Expansion of Sheared Holes with Burr In Contact With The Punch [Based on Data from Reference 11]](https://ahssinsights.org/wp-content/uploads/2020/07/Pasted-into-Hole-Expansion-Testing-Intro-32.jpg)
Figure 14: Effect of Punch Type on Hole Expansion of Sheared Holes with Burr In Contact With The Punch. (Based on Data from Citation P-13.)
![Figure 15: Effect of Punch Type on Hole Expansion of Sheared Holes with Burr Facing Away From The Punch [Based on Data from Reference 11]](https://ahssinsights.org/wp-content/uploads/2020/07/Pasted-into-Hole-Expansion-Testing-Intro-33.jpg)
Figure 15: Effect of Punch Type on Hole Expansion of Sheared Holes with Burr Facing Away From The Punch. (Based on Data from Citation P-13.)
![Figure 16: Effect of Punch Type on Hole Expansion of Reamed Holes [Based on Data from Reference 11]](https://ahssinsights.org/wp-content/uploads/2020/07/Pasted-into-Hole-Expansion-Testing-Intro-34.jpg)
Figure 16: Effect of Punch Type on Hole Expansion of Reamed Holes. (Based on Data from Citation P-13.)
Figure 17 compares the simulation results from expanding a perfect edge (no burr, no strain) with a conical punch on the left and a spherical punch on the right.W-2 The color scale, based on a “damage” parameter, shows that a spherical punch results in a more uniform distribution of damage, especially at the edge. This suggests that the impact of burr orientation on hole expansion is less significant for this punch geometry.
Flanging with a conical punch causes high circumferential strain and high damage values at the outer edge. The inner edge of the sheet initially presses against the punch, and later stretches during flanging. Since cracks initiate at the fracture zone, using a conical punch with the burr facing the punch leads to a greater hole expansion capability than when having the burr in contact with a spherical punch.
Fracture initiates at the edge, and orienting the burr so that it is in contact with the punch leads to a greater hole expansion value.
Figure 17: Distribution of damage values in simulated hole expansion tests conducted with a conical punch (left image) and a hemispherical punch (right image).W-2
Improving Hole Expansion with New Punch Shapes
As explained above, the degree to which a sheared edge can be stretched before fracture is a function of many parameters, including the shape of the punch. Also contributing is the hardness uniformity of the microstructural phases, where grades with components having high hardness differences are associated with relatively lower hole expansion capability.
Researchers evaluated the effects of punch design and clearance on hole expansion capability of dual phase and ferrite-bainite steels, each with a tensile strength of approximately 780 MPa.L-47
In addition to a conventional flat punch face, other punch types studied were those with a beveled face, a humped shape, and a newly designed punch which combines the benefits of the prior two types. A chamfered or beveled punch is known to reduce punch forces and reverse snap-through loads, while at the same time improve edge quality and hole expansion by minimizing the hardness increases found in the shear affected zone.
Use of the humped punch (Figure 18) led to hole expansion improvements of up to 10%, compared with a conventional flat punch when testing either the DP or FB products. When comparing the edge characteristics, the humped punch results in an increased rollover zone. The authors attributed this to the hump geometry imposing axial tension on the steel during punching, thereby increasing stress triaxiality. Increases in stress triaxiality results in a reduction in effective stress even at the same average stress. This in turn lowers the plastic deformation at the sheared edge which minimizes edge fracture. For these reasons, the increases in stress triaxiality associated with the humped punch promotes higher levels of hole expansion.
Figure 18: Humped punch design used in Citation L-47.
A newly designed punch which combines a beveled and humped design (Figure 19) increases hole expansion by more than 30% in both dual phase and ferrite-bainite steels (Figure 20). As explained above, the newly designed punch is effective in promoting stress triaxiality and minimizing the plastic deformation near the sheared edge. Furthermore, the beveled design improves the shear affected zone (SAZ) characteristics, leading to improved sheared edge expandability as measured in a hole expansion test.
Figure 19: New punch design incorporating features of beveled and humped punches.L-47
Figure 20: Effect of punch type and clearance on the hole expansion ratio of 780 MPa tensile strength steels. Left graph represents ferrite-bainite steel; right graph represents a dual phase steel. Legend: N=new punch design; C=conventional; H=humped; S=shear (beveled).L-47
Correlation of Hole Expansion Ratio with Tensile Properties
The complexities of hole expansion testing, as well as relatively few laboratories with the necessary test equipment and expertise, have led researchers to look for a correlation between the hole expansion ratio and conventionally measured properties obtained from a tensile test like the yield and tensile strength, uniform and total elongation, n-value, and r-value. Researchers even studied manipulations such as the yield-to-tensile ratio, tensile strength multiplied by uniform elongation, and n-value multiplied by r-value. Unfortunately, none of these properties or combinations have suitable correlation with the hole expansion ratio.
Recent work has shown a promising correlation between the hole expansion ratio and the true thinning strain at fracture. Our article on true fracture strain, describes this in greater detail.
![Figure 1: Transmission tunnel of 2005 Volkswagen Passat: (a) hot forming of pre-form, and final parts: (a) uncoated blank would suffer from scaling, (c) scale-free parts can be formed from varnish-coated blanks [REFERENCE 7]](http://ahssinsights.org/wp-content/uploads/2021/02/Fig1_Transmission-tunnel-of-2005-Volkswagen-Passat.jpg)
![Figure 2: Dry and wet areas in a car body. [REFERENCE 8]](http://ahssinsights.org/wp-content/uploads/2021/02/Fig2_Dry-and-wet-areas-in-a-car-body.jpg)
![Figure 3: Oxide layer (scale) on press hardened steel after: (a) fast resistance heating (10 seconds in air), (b) furnace heating (120 seconds in air) [REFERENCE 14]](http://ahssinsights.org/wp-content/uploads/2021/02/Fig3_Oxide-layer-on-press-hardened-steel-after-fast-resistance-heating.jpg)
![Figure 6: IDL thickness variation with furnace dwell time (Image created by REFERENCE 43 using raw data from REFERENCE 22, REFERENCE 26, and REFERENCE 27]](http://ahssinsights.org/wp-content/uploads/2021/02/Fig6_IDL-thickness-variation-with-furnace-dwell-time.jpg)
![Figure 10: Evolution of galvanized coating: (a) as delivered: Ferrite+Pearlite in base metal, almost pure Zn coating with Al-rich inhibition layer, (b) at high temperatures: austenite in base metal + α-Fe(Zn) and liquid Zn + surface oxides, (c) after press hardening: martensite in base metal + α-Fe(Zn) and Γ-phase coatings + surface oxides. The oxides are removed prior to welding and painting [REFERENCE 30]](http://ahssinsights.org/wp-content/uploads/2021/02/Fig10_Evolution-of-galvanized-coating.jpg)
![Figure 11: Manual application of a varnish coating. [REFERENCE 7]](http://ahssinsights.org/wp-content/uploads/2021/02/Fig11_Manual-application-of-a-varnish-coating.jpg)
