Die-Free Blanking of Class A Quality & Structural Parts

Die-Free Blanking of Class A Quality & Structural Parts

You’ll find this content as part of our page on Laser Blanking, but this month, we want to highlight it in our AHSS Insights blog.  We thank Schuler North America for contributing this insightful case study.

Production of Class A quality and structural parts without a blanking die is possible, even for high-volume serial production. Laser blanking enables flexible, cost-effective, and sustainable manufacturing and is capable of reaching 45 parts per minute.  DynamicFlow Technology (DFT) from Schuler provides highly productive, die-free blanking with lasers—directly from a continuously running steel coil. DFT combines the advantages of flexible laser cutting with the speed of conventional blanking.


Laser Blanking Lines at a Glance

Figure 1 – Laser blanking lines offer additional flexibility over conventional blanking approaches.

Laser blanking technology addresses market challenges such as frequent die changes, the need to increase capacity, and improving plant floor utilization, material utilization, and downstream processes.



It is important to remember that there are no dies with laser blanking technology, and no dies mean no die changes. Overall Equipment Effectiveness (OEE) of up to 80% can be achieved with laser blanking technology. In fact, 4 to 6 million parts per year of various materials are produced with the help of DFT—including mild steel, high-strength steel, and advanced high-strength steel. Even processing press-hardening steels with an aluminum-silicon coating is possible with laser blanking.  Surface and cutting quality can be maintained over this spectrum of steel grades.  Laser blanking technology can even achieve effective small batch production of Class A outer body panels and structural parts typically up to 3mm thick.



Competitive high-speed and high-output results can be achieved in multiple ways with laser blanking technology. The above-ground coil-fed line, optimized for short setup time, can handle coils with material widths up to 2,150 mm, weighing up to 30 tons. The material transport is smooth and controlled, simplifying setup and leading to uninterrupted processing within the laser cell.

There are three highly dynamic and simultaneously moving laser cutting heads within the laser cell of these lines. These laser cutting heads cut the programmed blank contour from a continuously moving material coil. Cutting speeds can exceed 100 meters per minute. The material is protected against any process contamination throughout the cutting process by custom-designed cutting clearance and material transport. 

Figure 2 reveals the high-speed and high-output results for outer body parts. Each part is measured by improved output per minute and hour to achieve an OEE of 80%. Laser blanking lines can achieve up to 45 parts per minute and reduce costs per blank.

High productivity achieved with laser blanking

Figure 2: High productivity achieved with laser blanking  



Up to 90% of blank costs are determined by the material price. The most significant leverage would be to reduce scrap and save on materials. Schuler conducted research based on the production of 300,000 cars per year, at 350 kg per car and $1,000 USD per ton of steel to provide a realistic inside look at how much cost savings can be achieved with laser blanking. The result was $1 Million USD saved with just 1% of material savings. This is extremely significant as material costs keep increasing.

Laser blanking is the digital way to cut blanks. All that’s needed to create a blanking program is a drawing to be loaded and a material to be selected. The part-specific program can be created offline and modified at any time. It is designed to create optimal combinations of material utilization and output—resulting in a high level of flexibility that significantly reduces development time for optimal blanks while also allowing for need-based production. This makes production planning easier, and it also opens the door to continuous contour optimizations for the forming process. Additionally, laser cutting does not require any gaps between individual parts due to smart nesting capabilities that cannot be achieved in comparison to die nesting or flatbed laser nesting. The combined smart, flexible nesting functions unlock new potential for material savings. Manufacturers can optimize individual blanks and eliminate the separating strip or connection bridges. Scrap savings in the forming process can also be achieved as there are no geometric restrictions due to cutting dies, and manufacturers can continuously optimize or adapt parts.

Figure 3 showcases the comparison of die nesting (the two graphics on the left) versus a laser-optimized blank contour and material savings via smart, laser blanking line nesting (the two images on the right). 

Die nesting compared with laser-optimized blank contours highlighting potential

Figure 3: Die nesting (left) compared with laser-optimized blank contours highlighting potential material savings (right)

Overall, laser blanking lines can have an equivalent throughput to conventional blanking lines, but laser blanking lines can achieve up to 10% greater material utilization.


You can read the full Case Study, including how laser blanking reduces infrastructure costs and improves downstream processes here: Laser Blanking Case Study


Schuler will present laser blanking technology, along with a variety of digital tools that create the “Press Shop of the Future” at FABTECH Chicago 2023 (booth # D41306). Tiago Vasconcellos, Sales Director at Schuler North America, will present “How Smart is Your Press Shop?” during FABTECH’s Educational Conference. The presentation will use The Smart Press Shop, a newly formed joint venture between Porsche and Schuler, as an exemplary case study for smart manufacturing standards. Attendees will discover innovative and practical ways to incorporate digitalization into production and become a state-of-the-art stamping facility directly from Schuler. 

About Schuler Group

Schuler offers customized cutting-edge technology in all areas of forming—from the networked press to press shop planning. In addition to presses, Schuler’s products include automation, dies, process know-how, and service for the entire metalworking industry. Schuler’s Digital Suite brings together solutions for networking forming technology and is continuously being developed to further improve line productivity and availability. Schuler customers include automotive manufacturers and suppliers, as well as companies in the forging, household appliance, and electrical industries. Schuler presses are minting coins for more than 180 countries. Founded in 1839 at the Göppingen, Germany headquarters, Schuler has approximately 5,000 employees at production sites in Europe, China and the Americas, as well as service companies in more than 40 countries. The company is part of the international technology group ANDRITZ.

Schuler’s global portfolio of world-renowned brands include BCN (Bliss Clearing Niagara) Technical Services, Müller Weingarten, Beutler, Umformtechnik Erfurt, SMG Pressen, Hydrap Pressen, Wilkins & Mitchell, Bêché, Spiertz Presses, Farina Presse, Liebergeld, Peltzer & Ehlers, Schleicher, and Sovema Group.

About Schuler North America

Schuler North America (Schuler), headquartered in Canton, Michigan, is the North American subsidiary of Schuler Group. Schuler provides new equipment, spare parts, and a portfolio of lifecycle services for all press systems—including preventative maintenance, press shop design and optimization, turnkey installations, retrofits for existing systems, and localized production and service. Schuler’s best-in-class position in the metalworking and materials industry serves automotive manufacturers and tier suppliers, as well as home appliance, electronics, forging, and other industries.


Roll Forming

Roll Forming

Roll Forming takes a flat sheet or strip and feeds it longitudinally through a mill containing several successive paired roller dies, each of which incrementally bends the strip into the desired final shape. The incremental approach can minimize strain localization and compensate for springback. Therefore, roll forming is well suited for generating many complex shapes from Advanced High-Strength Steels, especially from those grades with low total elongation, such as martensitic steel. The following video, kindly provided by Shape Corp.S-104, highlights the process that can produce either open or closed (tubular) sections.

The number of pairs of rolls depends on the sheet metal grade, finished part complexity, and the design of the roll-forming mill. A roll-forming mill used for bumpers may have as many as 30 pairs of roller dies mounted on individually driven horizontal shafts.A-32

Roll forming is one of the few sheet metal forming processes requiring only one primary mode of deformation. Unlike most forming operations, which have various combinations of forming modes, the roll-forming process is nothing more than a carefully engineered series of bends. In roll forming, metal thickness does not change appreciably except for a slight thinning at the bend radii.

Roll forming is appropriate for applications requiring high-volume production of long lengths of complex sections held to tight dimensional tolerances. The continuous process involves coil feeding, roll forming, and cutting to length. Notching, slotting, punching, embossing, and curving combine with contour roll forming to produce finished parts off the exit end of the roll-forming mill. In fact, companies directly roll-form automotive door beam impact bars to the appropriate sweep and only need to weld on mounting brackets prior to shipment to the vehicle assembly line.A-32 Figure 1 shows an example of automotive applications that are ideal for the roll-forming process.

Figure 1: Body components that are ideally suited for roll-forming.

Figure 1: Body components that are ideally suited for roll-forming.

Roll forming can produce AHSS parts with:

  • Steels of all levels of mechanical properties and different microstructures.
  • Small radii depending on the thickness and mechanical properties of the steel.
  • Reduced number of forming stations compared with lower strength steel.

However, the high sheet-steel strength means that forces on the rollers and frames in the roll-forming mill are higher. A rule of thumb says that the force is proportional to the strength and thickness squared. Therefore, structural strength ratings of the roll forming equipment must be checked to avoid bending of the shafts. The value of minimum internal radius of a roll formed component depends primarily on the thickness and the tensile strength of the steel (Figure 2).

Figure 2: Achievable minimum r/t values for bending and roll forming for different strength and types of steel.S-5

Figure 2: Achievable minimum r/t values for bending and roll forming for different strength and types of steel.S-5

As seen in Figure 2, roll forming allows smaller radii than a bending process. Figure 3 compares CR1150/1400-MS formed with air-bending and roll forming. Bending requires a minimum 3T radius, but roll forming can produce 1T bends.S-30

Figure 3: CR1150/1400-MS (2 mm thick) has a minimum bend radius of 3T, but can be roll formed to a 1T radius.S-30

Figure 3: CR1150/1400-MS (2 mm thick) has a minimum bend radius of 3T, but can be roll formed to a 1T radius.S-30

The main parameters having an influence on the springback are the radius of the component, the sheet thickness, and the strength of the steel. As expected, angular change increases for increased tensile strength and bend radius (Figure 4).

Figure 4: Angular change increases with increasing tensile strength and bend radii.A-4

Figure 4: Angular change increases with increasing tensile strength and bend radii.A-4

Figure 5 shows a profile made with the same tool setup for three steels at the same thickness having tensile strength ranging from 1000 MPa to 1400 MPa. Even with the large difference in strength, the springback is almost the same.

Figure 5: Roll formed profile made with the same tool setup for three different steels. Bottom to Top: CR700/1000-DP, CR950/1200-MS, CR1150/1400-MS.S-5

Figure 5: Roll formed profile made with the same tool setup for three different steels. Bottom to Top: CR700/1000-DP, CR950/1200-MS, CR1150/1400-MS.S-5

Citation A-33 provides guidelines for roll forming High-Strength Steels:

  • Select the appropriate number of roll stands for the material being formed. Remember the higher the steel strength, the greater the number of stands required on the roll former.
  • Use the minimum allowable bend radius for the material in order to minimize springback.
  • Position holes away from the bend radius to help achieve desired tolerances.
  • Establish mechanical and dimensional tolerances for successful part production.
  • Use appropriate lubrication.
  • Use a suitable maintenance schedule for the roll forming line.
  • Anticipate end flare (a form of springback). End flare is caused by stresses that build up during the roll forming process.
  • Recognize that as a part is being swept (or reformed after roll forming), the compression of metal can cause sidewall buckling, which leads to fit-up problems.
  • Do not roll form with worn tooling, as the use of worn tools increases the severity of buckling.
  • Do not expect steels of similar yield strength from different steel sources to behave similarly.
  • Do not over-specify tolerances.

Guidelines specifically for the highest strength steelsA-33:

  • Depending on the grade, the minimum bend radius should be three to four times the thickness of the steel to avoid fracture.
  • Springback magnitude can range from ten degrees for 120X steel (120 ksi or 830 MPa minimum yield strength, 860 MPa minimum tensile strength) to 30 degrees for M220HT (CR1200/1500-MS) steel, as compared to one to three degrees for mild steel. Springback should be accounted for when designing the roll forming process.
  • Due to the higher springback, it is difficult to achieve reasonable tolerances on sections with large radii (radii greater than 20 times the thickness of the steel).
  • Rolls should be designed with a constant radius and an evenly distributed overbend from pass to pass.
  • About 50 percent more passes (compared to mild steel) are required when roll forming ultra high-strength steel. The number of passes required is affected by the number of profile bends, mechanical properties of the steel, section depth-to-steel thickness ratio, tolerance requirements, pre-punched holes and notches.
  • Due to the higher number of passes and higher material strength, the horsepower requirement for forming is increased.
  • Due to the higher material strength, the forming pressure is also higher. Larger shaft diameters should be considered. Thin, slender rolls should be avoided.
  • During roll forming, avoid undue permanent elongation of portions of the cross section that will be compressed during the sweeping process.

Roll forming is applicable to shapes other than long, narrow parts. For example, an automaker roll forms their pickup truck beds allowing them to minimize thinning and improve durability (Figure 6). Reduced press forces are another factor that can influence whether a company roll forms rather than stamps truck beds.

Figure 6: Roll Forming can replace stamping in certain applications.G-9

Figure 6: Roll Forming can replace stamping in certain applications.G-9

Traditional two-dimensional roll forming uses sequential roll stands to incrementally change flat sheets into the targeted shape having a consistent profile down the length. Advanced dynamic roll forming incorporates computer-controlled roll stands with multiple degrees of freedom that allow the finished profile to vary along its length, creating a three-dimensional profile. The same set of tools create different profiles by changing the position and movements of individual roll stands. In-line 3D profiling expands the number of applications where roll forming is a viable parts production option.

One such example are the 3D roll formed tubes made from 1700 MPa martensitic steel for A-pillar / roof rail applications in the 2020 Ford Explorer and 2020 Ford Escape (Figure 7).  Using this approach instead of hydroforming created smaller profiles resulting in improved driver visibility, more interior space, and better packaging of airbags. The strength-to-weight ratio improved by more than 50 percent, which led to an overall mass reduction of 2.8 to 4.5 kg per vehicle.S-104

Figure 7: 3D Roll Formed Profiles in 2020 Ford Vehicles using 1700 MPa martensitic steel.S-104

Figure 7: 3D Roll Formed Profiles in 2020 Ford Vehicles using 1700 MPa martensitic steel.S-104

Roll forming is no longer limited to producing simple circular, oval, or rectangular profiles. Advanced cross sections such
as those shown in Figure 8 provided by Shape Corporation highlight some profile designs aiding in body structure
stiffness and packaging space reductions.

Roll Form Designs

Figure 8: Roll forming profile design possibilities. Courtesy of Shape Corporation.

In summary, roll forming can produce AHSS parts with steels of all levels of mechanical properties and different microstructures with a reduced R/T ratio versus conventional bending. All deformation occurs at a radius, so there is no sidewall curl risk and overbending works to control angular springback.

Case Study: How Steel Properties Influence the Roll Forming Process

Many thanks to Brian Oxley, Product Manager, Shape Corporation, and Dr. Daniel Schaeffler, President, Engineering
Quality Solutions, Inc., for providing this case study.

Optimizing the use of roll forming requires understanding how the sheet metal behaves through the process.
Making a bend in a roll formed part occurs only when forming forces exceed the metal’s yield strength, causing plastic
deformation to occur. Higher strength sheet metals increase forming force requirements, leading to the need to have
larger shaft diameters in the roll forming mill. Each pass must have greater overbend to compensate for the increasing
springback associated with the higher strength.

Figure 9 provides a comparison of the loads on each pass of a 10-station roll forming line when forming either AISI 1020
steel (yield strength of 350 MPa, tensile strength of 450 MPa, elongation to fracture of 15%) or CR1220Y1500T-MS, a
martensitic steel with 1220 MPa minimum yield strength and 1500 MPa minimum tensile strength.


Graph showing pass loads with maximum force levels. Gold and purple 3D bar chart

Figure 9: Loads on each pass of a roll forming line when forming either AISI 1020 steel (450 MPa tensile
strength) or a martensitic steel with 1500 MPa minimum tensile strength. Courtesy of Roll-Kraft.


Although a high-strength material requires greater forming loads, grades with higher yield strength can resist stretching
of the strip edge and prevent longitudinal deformations such as twisting or bow. Flange edge flatness after forming
either AISI 1020 or CR1220Y1500T-MS is presented in Figure 10.


Figure 10: Simulation results showing flange edge flatness of a) AISI 1020 and b) CR1220Y1500T-MS.
Assumptions for the simulation: AISI 1020 yield strength = 350 MPa; CR1220Y1500T yield strength = 1220 MPa. Higher yield strength leads to better flatness.


Force requirements for piercing operations are a function of the sheet tensile strength. High strains in the part design
exceeding uniform elongation resulting from loads in excess of the tensile strength produces local necking, representing
a structural weak point. However, assuming the design does not produce these high strains, the tensile strength has only
an indirect influence on the roll forming characteristics.

Yield strength and flow stress are the most critical steel characteristics for roll forming dimensional control. Receiving
metal with limited yield strength variability results in consistent part dimensions and stable locations for pre-pierced

Flow stress represents the strength after some amount of deformation, and is therefore directly related to the degree of
work hardening: starting at the same yield strength, a higher work hardening steel will have a higher flow stress at the
same deformation.

Two grades are shown in Figure 11: ZE 550 and CR420Y780T-DP. ZE 500, represented by the red curve, is a recovery
annealed grade made by Bilstein having a yield strength range of 550 to 625 MPa and a minimum tensile strength of 600
MPa, while CR420Y780T-DP, represented by the blue curve, is a conventional dual phase steel with a minimum yield
strength of 420 MPa and a minimum tensile strength of 780 MPa. For the samples tested, ZE 550 has a yield strength of
approximately 565 MPa, where that for CR420Y780T-DP is much lower at about 485 MPa. Due to the higher work
hardening (n-value) of the DP steel, its flow stress at 5% strain is 775 MPa, while the flow stress for the HSLA grade at 5%
strain is 620 MPa.

In conventional stamping operations, this work hardening is beneficial to delay the onset of necking. However, use of
dual-phase steels and other grades with high n-value can lead to dimensional issues in roll-formed parts. Flow stress in a
given area is a function of the local strain. Each roll station induces additional strain on the overall part, and strains vary
within the part and along the edge. This strength variation is responsible for differing springback and edge wave across
a roll-formed part.

Unlike conventional stamping, grades with a high yield/tensile ratio where the yield strength is close to the tensile
strength are better suited to produce straight parts via roll forming.

Line graph of stress/strain curves

Figure 11: Stress-strain curves for CR420Y780T-DP (blue) and ZE 550 (red). See text for description of the grades.


Total elongation to fracture is the strain at which the steel breaks during tensile testing, and is a value commonly
reported on certified metal property documents (cert sheets). As observed on the colloquially called “banana diagram”,
elongation generally decreases as the strength of the steel increases.

For lower strength steels, total elongation is a good indicator for a metal’s bendability. Bend severity is described by the
r/t ratio, or the ratio of the inner bend radius to the sheet thickness. The metal’s ability to withstand a given bend can be
approximated by the tensile test elongation, since during a bend, the outermost fibers elongate like a tensile test.

In higher strength steels where the phase balance between martensite, bainite, austenite, and ferrite play a much larger
role in developing the strength and ductility than in other steels, bendability is usually limited by microstructural
uniformity. Dual phase steels, for example, have excellent uniform elongation and resistance to necking coming from
the hardness difference between ferrite and martensite. However, this large hardness difference is also responsible for
relatively poor edge stretchability and bendability. In roll forming applications, those grades with a uniform
microstructure will typically have superior performance. As an example, refer to Figure 11. The dual phase steel shown
in blue can be bent to a 2T radius before cracking, but the recovery annealed ZE 550 grade with noticeably higher yield
strength and lower elongation can be bent to a ½T radius.

Remember that each roll forming station only incrementally deforms the sheet, with subsequent stations working on a
different region. Roll formed parts do not need to use grades associated with high total elongation, especially since
these typically have a bigger gap between yield and tensile strength.


Coil Shape Imperfections Influencing Roll Forming

Along with the mechanical properties of steel, physical shape attributes of the sheet or coil can influence the roll
forming process. These include center buckle, coil set, cross bow, and camber. Receiving coils with these imperfections
may result in substandard roll formed parts.
Flatness is paramount when it comes to getting good shape on roll formed parts. Individual OEMs or processors may
have company-specific procedures and requirements, while organizations like ASTM offer similar information in the
public domain. ASTM A1030/A1030M is one standard covering the practices for measuring flatness, and specification
ASTM A568/A568M shows methods for characterizing longitudinal waves, buckles, and camber. (link to
https://www.astm.org/a1030_a1030m-21.html and https://www.astm.org/a0568_a0568m-19a.html)

Center buckle (Figure 12), also known as full center, is the term to describe pockets or waves in the center or quarter
line of the strip. The height of pocket varies from 1/6” to 3/4”. Center buckle occurs when the central width portion of
the master coil is longer than the edges. This over-rolling of the center portion might occur when there is excessive
crown in the work roll, build-up from the hot strip mill, a mismatched set of work rolls, improper use of the benders, or
improper rolling procedures. A related issue is edge buckle presenting as wavy edges, originating when the coil edges
are longer than the central width position.


Center Buckle

Figure 12: Coil shape imperfection: Center Buckle


Coil set (Figure 13a), also known as longitudinal bow, occurs when the top surface of the strip is stretched more than the
bottom surface, causing a bow condition parallel with the rolling direction. Here, the strip exhibits a tendency to curl
rather than laying flat. To some extent, coil set is normal, and easy to address with a leveler. Severe coil set may be
induced by an imbalance in the stresses induced during rolling by the thickness reduction work rolls. Potential causes include different diameters or surface speeds of the two work rolls, or different frictional conditions along the two arcs
of contact.

Crossbow (Figure 13b) is a bow condition perpendicular to the rolling direction, and arcs downward from the high point
in the center position across the width of the sheet. Crossbow may occur if improper coil set correction practices are

Figure 13: Coil shape imperfections: A) Coil set and B) Crossbow

Figure 13 caption: Coil shape imperfections: A) Coil set and B) Crossbow A-30

Camber (Figure 14) is the deviation of a side edge from a straight edge, and results when one edge of the steel is
elongated more than the other during the rolling process due to a difference in roll diameter or speed. The maximum
allowable camber under certain conditions is contained within specification ASTM A568/A568M, among others.

Figure 14 - Camber

Figure 14: Coil shape imperfection: Camber


Coil shape imperfections produce residual stresses in the starting material. These residual stresses combined with the
stresses from forming lead to longitudinal deviations from targeted dimensions after roll forming. Some of the resultant
shapes of roll formed components made from coils having these issues are shown in Figure 15. Leveling the coil prior to
roll forming may address some of these shape concerns, and has the benefit of increasing the yield strength, making a
more uniform product.


Figure 15

Figure 15: Shape deviations in roll formed components initiating from incoming coil shape issues: a) camber b) longitudinal bow c) twist d) flare e) center wave (center buckle) f) edge wave. H-66


Roll Stamping

Traditional roll forming creates products with essentially uniform cross sections.  A newer technique called Roll Stamping enhances the ability to create shapes and features which are not in the rolling axis.

Using a patented processA-48, R-9, forming rolls with the part shape along the circumferential direction creates the desired form, as shown in Figure 16.

Figure 7: Roll Stamping creates additional shapes and features beyond capabilities of traditional roll forming. (Reference 1)

Figure 16: Roll Stamping creates additional shapes and features beyond capabilities of traditional roll forming. A-48

This approach can be applied to a conventional roll forming line.  In the example of an automotive door impact beam, the W-shaped profile in the central section and the flat section which attaches to the door inner panel are formed at the same time, without the need for brackets or internal spot welds (Figure 17. Sharp corner curvatures are possible due to the incremental bending deformation inherent in the process.

Figure 8: A roll stamped door beam formed on a conventional roll forming line eliminates the need for welding brackets at the edges. (Reference 2)

Figure 17: A roll stamped door part formed on a conventional roll forming line eliminates the need for welding brackets at the edges.R-9

A global automaker used this method to replace a three-piece door impact beam made with a 2.0 mm PHS-CR1500T-MB press hardened steel tube requiring 2 end brackets formed from 1.4 mm CR-500Y780T-DP to attach it to the door frame, shown in Figure 18. The new approach, with a one-piece roll stamped 1.0 mm CR900Y1180T-CP complex phase steel impact beam, resulted in a 10% weight savings and 20% cost savings.K-58 This technique started in mass production on a Korean sedan in 2017, a Korean SUV in 2020, and a European SUV in 2021.K-58


Figure 10: Some Roll Stamping Automotive Applications (Citation D)

Figure 18: Some Roll Stamping Automotive Applications.K-58


Photo of Brian OxleyThanks are given to Brian Oxley, Product Manager, Shape Corporation, for his contributions to the Roll Forming Case Study and Coil Shape Imperfections section. Brian Oxley is a Product Manager in the Core Engineering team at Shape Corp. Shape Corp. is a global, full-service supplier of lightweight steel, aluminum, plastic, composite and hybrid engineered solutions for the automotive industry. Brian leads a team responsible for developing next generation products and materials in the upper body and closures space that complement Shape’s core competency in roll forming. Brian has a Bachelor of Science degree in Material Science and Engineering from Michigan State University.

Uniform Elongation

Uniform Elongation

During a tensile test, the elongating sample leads to a reduction in the cross-sectional width and thickness. The shape of the engineering stress-strain curve showing a peak at the load maximum (Figure 1) results from the balance of the work hardening which occurs as metals deform and the reduction in cross-sectional width and thickness which occurs as the sample dogbone is pulled in tension. In the upward sloping region at the beginning of the curve, the effects of work hardening dominate over the cross-sectional reduction. Starting at the load maximum (ultimate tensile strength), the reduction in cross-sectional area of the test sample overpowers the work hardening and the slope of the engineering stress-strain curve decreases. Also beginning at the load maximum, a diffuse neck forms usually in the middle of the sample.

Figure 1: Engineering stress-strain curve from which mechanical properties are derived.

Figure 1: Engineering stress-strain curve from which mechanical properties are derived.


The elongation at which the load maximum occurs is known as Uniform Elongation. In a tensile test, uniform elongation is the percentage the gauge length elongated at peak load relative to the initial gauge length. For example, if the gauge length at peak load measures 61 mm and the initial gauge length was 50mm, uniform elongation is (61-50)/50 = 22%.

Schematics of tensile bar shapes are shown within Figure 1. Note the gauge region highlighted in blue. Up though uniform elongation, the cross-section has a rectangular shape. Necking begins at uniform elongation, and the cross section is no longer rectangular.

Theory and experiments have shown that uniform elongation expressed in true strain units is numerically equivalent to the instantaneous n-value.

Deformation Prior to Uniform Elongation is Not Uniformly Distributed

Conventional wisdom for decades held that there is a uniform distribution of strains within the gauge region of a tensile bar prior to strains reaching uniform elongation. Traditional extensometers calibrated for 50-mm or 80-mm gauge lengths determine elongation from deformation measured relative to this initial length. This approach averages results over these spans.

The advent of Digital Image Correlation (DIC) and advanced processing techniques allowed for a closer look. A studyS-113 released in 2021 clearly showed that each of the 201 data points monitored within a 50 mm gauge length (virtual gauge length of 0.5-mm) experiences a unique strain evolution, with differences starting before uniform elongation.

Figure 2 caption: Strain evolution of the 201 points on the DP980 tensile-test specimen exhibits divergence beginning before uniform elongation—counter to conventional thinking.

Figure 2: Strain evolution of the 201 points on the DP980 tensile-test specimen exhibits divergence beginning before uniform elongation—counter to conventional thinking.S-113


High Strain Rate Testing

High Strain Rate Testing

Dynamic tensile testing of sheet steels is becoming more important due to the need for more optimized vehicle crashworthiness analysis in the automotive industry. Positive strain rate sensitivity (strength increases with strain rate) as an example, offers a potential for improved energy absorption during a crash event. New systems have been developed in recent years to meet the increasing demand for dynamic testing.

Three important points collectively highlight the need for high-speed testing:

  • Tensile properties and fracture behavior change with strain rate.
  • Conventional tensile tests using standard dogbone shapes take on the order of 1 to 2 minutes depending on the grade.
  • An entire automotive crash takes on the order of 100 milliseconds, with deformation rates 10,000 to 100,000 faster than conventional testing speeds.

Characterizing the response during high-speed testing provides critical information used in crash simulations, but these tests often require upgraded equipment and procedures. Conventional tensile testing equipment may lack the ability to reach the required speeds (on the order of 20 m/s). Sensors for load and displacement must acquire accurate data during tests which take just a few milliseconds.

Higher speed tensile and fracture characterization also aids in predicting the properties of stamped parts, as deformation rates in stamping are 100 to 1,000 times higher than most testing rates.

Steel alloys possess positive strain rate sensitivity, or m-value, meaning that strength increases with strain rate. This has benefits related to improved crash energy absorption.

Characterizing this response requires use of robust testing equipment and practices appropriate for the targeted strain rate. Some techniques involve a tensile or compressive Split Hopkinson (Kolsky) Bar, a drop tower or impact system, or a high-speed servo-hydraulic system. Historically, no guidelines were available as to the testing method, specimen dimensions, measurement devices, and other important issues which are critical to the quality of testing results. As a result, data from different laboratories were often not comparable. A WorldAutoSteel committee evaluated various procedures, conducted several round-robins, and developed a recommended procedure, which evolved into what are now both parts of ISO 26203, linked below.

Published standards addressing tensile testing at high strain rates include:

The specific response as a function of strain rate is grade dependent. Some grades get stronger and more ductile as the strain rate increases (left image in Figure 1), while other grades see primarily a strength increase (right image in Figure 1). Increases are not linear or consistent with strain rate, so simply scaling the response from conventional quasi-static testing does not work well. Strain hardening (n-value) also changes with speed in some grades, as suggested by the different slopes in the right image of Figure 1. Accurate crash models must also consider how strain rate sensitivity impacts bake hardenability and the magnitude of the TRIP effect, both of which are further complicated by the strain levels in the part from stamping.

Figure 1: Two steels with different strength/ductility response to increasing strain rate.A-7

Figure 1: Two steels with different strength/ductility response to increasing strain rate.A-7




Simulation Inputs

Simulation Inputs


Predicting metal flow and failure is the essence of sheet metal forming simulation.  Characterizing the stress-strain response to metal flow requires a detailed understanding of when the sheet metal first starts to permanently deform (known as the yield criteria), how the metal strengthens with deformation (the hardening law), and the failure criteria (for example, the forming limit curve). Complicating matters is that each of these responses changes as three-dimensional metal flow occurs, and are functions of temperature and forming speed. 

The ability to simulate these features reliably and accurately requires mathematical constitutive laws that are appropriate for the material and forming environments encountered. Advanced models typically improve prediction accuracy, at the cost of additional numerical computational time and the cost of experimental testing to determine the material constants. Minimizing these costs requires compromises, with some of these indicated in Table I created based on Citations B-16 and R-28.

Table I: Deviations from reality made to reduce simulation costs. Based on Citations B-16 and R-28.

Table I: Deviations from reality made to reduce simulation costs. Based on Citations B-16 and R-28.


Yield Criteria

The yield criteria (also known as the yield surface or yield loci) defines the conditions representing the transition from elastic to plastic deformation.  Assuming uniform metal properties in all directions allows for the use of isotropic yield functions like von Mises or Tresca. A more realistic approach considers anisotropic metal flow behavior, requiring the use of more complex yield functions like those associated with Hill, Barlat, Banabic, or Vegter.   

No one yield function is best suited to characterize all metals. Some yield functions have many required inputs.  For example, “Barlat 2004-18p” has 18 separate parameters leading to improved modeling accuracy – but only when inserting the correct values. Using generic textbook values is easier, but negates the value of the chosen model.  However, determining these variables typically is costly and time-consuming, and requires the use of specialized test equipment.

Hardening Curve

Metals get stronger as they deform, which leads to the term work hardening. The flow stress at any given amount of plastic strain combines the yield strength and the strengthening from work hardening.  In its simplest form, the stress-strain curve from a uniaxial tensile test shows the work hardening of the chosen sheet metal. This approach ignores many of the realities occurring during forming of engineered parts, including bi-directional deformation.

Among the simpler descriptions of flow stress are those from Hollomon, Swift, and Ludvik.  More complex hardening laws are associated with Voce and Hockett-Sherby. 

The strain path followed by the sheet metal influences the hardening. Approaches taken in the Yoshida-Uemori (YU) and the Homogeneous Anisotropic Hardening (HAH) models extend these hardening laws to account for Bauschinger Effect deformations (the bending-unbending associated with travel over beads, radii, and draw walls).

As with the yield criteria, accuracy improves when accounting for three-dimensional metal flow, temperature, and forming speed, and using experimentally determined input parameters for the metal in question rather than generic textbook values. 

Failure Conditions

Defining the failure conditions is the other significant challenge in metal forming simulation. Conventional Forming Limit Curves describe necking failure under certain forming modes, and are easier to understand and apply than alternatives. Complexity and accuracy increase when accounting for non-linear strain paths using stress-based Forming Limit Curves.  Necking failure is not the only type of failure mode encountered. Conventional FLCs cannot predict fracture on tight radii and cut edges, nor can they account for dimensional issues like springback.  For these, failure criteria definitions which are more mathematically complex are appropriate.

Constitutive Laws and Their Influence

on Forming Simulation Accuracy

Many simulation packages allow for an easy selection of constitutive laws, typically through a drop-down menu listing all the built-in choices. This ease potentially translates into applying inappropriate selections unless the simulation analyst has a fundamental understanding of the options, the inputs, and the data generation procedures.

Some examples:

  • The “Keeler Equation” for the estimation of FLC0 has many decades of evidence in being sufficiently accurate when applied to mild steels and conventional high strength steels. The simple inputs of n-value and thickness make this approach particularly attractive.  However, there is ample evidence that using this approach with most advanced high strength steels cannot yield a satisfactory representation of the Forming Limit Curve.
  • Even in cases where it is appropriate to use the Keeler Equation, a key input is the n-value or the strain hardening exponent. This value is calculated as the slope of the (natural logarithm of the true stress):(natural logarithm of the true strain curve). The strain range over which this calculation is made influences the generated n-value, which in turn impacts the calculated value for FLC0.
  • The strain history as measured by the strain path at each location greatly influences the Forming Limit. However, this concept has not gained widespread understanding and use by simulation analysts.
  • A common method to experimentally determine flow curves combines tensile testing results through uniform elongation with higher strain data obtained from biaxial bulge testing. Figure 1 shows a flow curve obtained in this manner for a bake hardenable steel with 220 MPa minimum yield strength.  Shown in Figure 2 is a comparison of the stress-strain response from multiple hardening laws associated with this data, all generated from the same fitting strain range between yield and tensile strength.  Data diverges after uniform elongation, leading to vastly different predictions. Note that the differences between models change depending on the metal grade and the input data, so it is not possible to say that one hardening law will always be more accurate than others.

Figure 1: Flow curves for a bake hardenable steel generated by combinng tensile testing with bulge testing L-20

Figure 1: Flow curves for a bake hardenable steel generated by combining tensile testing with bulge testing.L-20


Figure 2: The chosen hardening law leads to vastly different predictions of stress-strain responses L-20

Figure 2: The chosen hardening law leads to vastly different predictions of stress-strain responses.L-20


  • Analysts often treat Poisson’s Ratio and the Elastic Modulus as constants.  It is well known that the Bauschinger Effect leads to changes in the Elastic Modulus, and therefore impacts springback.  However, there are also significant effects in both Poisson’s Ratio (Figure 3) and the Elastic Modulus (Figure 4) as a function of orientation relative to the rolling direction. Complicating matters is that this effect changes based on the selected metal grade.  

Figure 3:  Poisson’s Ratio as a Function of Orientation for Several Grades (Drawing Steel, DP 590, DP 780, DP 1180, and MS 1700) D-11

Figure 3:  Poisson’s Ratio as a Function of Orientation for Several Grades (Drawing Steel, DP 590, DP 980, DP 1180, and MS 1700) D-11


Figure 4:  Modulus of Elasticity as a Function of Orientation for Several Grades (Drawing Steel, DP 590, DP 780, DP 1180, and MS 1700) D-11

Figure 4:  Modulus of Elasticity as a Function of Orientation for Several Grades (Drawing Steel, DP 590, DP 980, DP 1180, and MS 1700) D-11


Testing to Determine Inputs for Simulation

Complete material card development requires results from many tests, each attempting to replicate one or more aspects of metal flow and failure. Certain models require data from only some of these tests, and no one model typically is best for all metals and forming conditions.  Tests described below include:

  • Tensile testing [room temperature at slow strain rates to elevated temperature with accelerated strain rates]
  • Biaxial bulge testing
  • Biaxial tensile testing
  • Shear testing
  • V-bending testing
  • Tension-compression testing with cyclic loading
  • Friction

Tensile testing is the easiest and most widely available mechanical property evaluation required to generate useful data for metal forming simulation. However, a tensile test provides a complete characterization of material flow only when the engineered part looks like a dogbone and all deformation resulted from pulling the sample in tension from the ends. That is obviously not realistic. Getting tensile test results in more than just the rolling direction helps, but generating those still involves pulling the sample in tension.  Three-dimensional metal flow occurs, and the stress-strain response of the sheet metal changes accordingly.  

The uniaxial tensile test generates a draw deformation strain state since the edges are free to contract.  A plane strain tensile test requires using a modified sample geometry with an increased width and decreased gauge length, 

Forming all steels involves a thermal component, either resulting from friction and deformation during “room temperature” forming or the intentional addition of heat such as used in press hardening. In either case, modeling the response to temperature requires data from tests occurring at the temperature of interest, at appropriate forming speeds.  Thermo-mechanical simulators like Gleeble™ generate such data.

Conventional tensile testing occurs at deformation rates of 0.001/sec. Most production stamping occurs at 10,000x that amount, or 10/sec. Crash events can be 2 orders of magnitude faster, at about 1000/sec.  The stress-strain response varies by both testing speed and grade. Therefore, accurate simulation models require data from higher-speed tensile testing. Typically, generating high speed tensile data involves drop towers or Split Hopkinson Pressure Bars.

A pure uniaxial stress state exists in a tensile test only until reaching uniform elongation and the beginning of necking.  Extrapolating uniaxial tensile data beyond uniform elongation risks introducing inaccuracies in metal flow simulations. Biaxial bulge testing generates the data for yield curve extrapolation beyond uniform elongation. This stretch-forming process deforms the sheet sample into a dome shape using hydraulic pressure, typically exerted by water-based fluids.  Citation I-12 describes a standard test procedure for biaxial bulge testing.

A Marciniak test used to create Forming Limit Curves generates in-plane biaxial strains.  Whereas FLC generation uses 100 mm diameter samples, larger samples allow for extraction of full-size tensile bars.  Although this approach generates samples containing biaxial strains, the extracted samples are tested uniaxially in the conventional manner.

Biaxial tensile testing allows for the determination of the yield locus and the biaxial anisotropy coefficient, which describes the slope of the yield surface at the equi-biaxial stress state. This test uses cruciform-shaped test pieces with parallel slits cut into each arm. Citation I-13 describes a standard test procedure for biaxial tensile testing.  The biaxial anisotropy coefficient can also be determined using the disk compression testing as described in Citation T-21.

Shear testing characterizes the sheet metal in a shear loading condition. There is no consensus on the specimen type or testing method. However, the chosen testing set-up should avoid necking, buckling, and any influence of friction.

V-bending tests determine the strain to fracture under specific loading conditions. Achieving plane strain or plane stress loading requires use of a test sample with features promoting the targeted strain state. 

Tension-compression testing characterizes the Bauschinger Effect.  Multiple cycles of tension-compression loading captures cyclic hardening behavior and elastic modulus decay, both of which improve the accuracy of springback predictions.  Again, no standard procedure exists. The biggest challenge with this test is preventing buckling from occurring during in-plane compressive loading. Related to this is the need to compensate for the friction caused by the anti-buckling mechanism in the stress-strain curves .

Friction is obviously a key factor in how metal flows.  However, there is no one simple value of friction that applies to all surfaces, lubricants, and tooling profiles. The coefficient of friction not only varies from point to point on each stamping but changes during the forming process. Determining the coefficient of friction experimentally is a function of the testing approach used. The method by which analysts incorporate friction into simulations influences the accuracy and applicability of the results of the generated model.

Studies are underway to reduce the costs and challenges of obtaining much of this data. It may be possible, for example, to use Digital Image Correlation (DIC) during a simple uniaxial tensile testing to quantify r-value at high strains, determine the material hardening behavior along with strain rate sensitivity, assess the degradation of Young’s Modulus during unloading, and use the detection of the onset of local neck to help account for non-linear strain path effects.S-110


Application of Advanced Testing to Failure Predictions

Global formability failures occur when the forming strains exceed the necking forming limit throughout the entire thickness of the sheet. Advanced steels are at risk of local formability failures where the forming strains exceed the fracture forming limit at any portion of the thickness of the sheet.

Fracture forming limit curves plot higher than the conventional necking forming limit curves on a graph showing major strain on the vertical axis and minor strain on the horizontal axis.  In conventional steels the gap between the fracture FLC and necking FLC is relatively large, so the part failure is almost always necking.  The forming strains are not high enough to reach the fracture FLC.

In contrast, AHSS grades are characterized by a smaller gap between the necking FLC and the fracture FLC.  Depending on the forming history, part geometry (tight radii), and blank processing (cut edge quality), forming strains may exceed the fracture FLC at an edge or bend before exceeding the necking FLC through-thickness.  In this scenario, the part will fracture without signs of localized necking.

A multi-year study funded by the American Iron and Steel Institute at the University of Waterloo Forming and Crash Lab describes a methodology used for forming and fracture characterization of advanced high strength steels, the details of which can be found in Citations B-11, W-20, B-12, B-13, R-5, N-13 and G-19.

This collection of studies, as well as work coming out of these studies, show that relatively few tests sufficiently characterize forming and fracture of AHSS grades.  These studies considered two 3rd Gen Steels, one with 980MPa tensile strength and one with 1180MPa tensile.

  • The yield surface as generated with the Barlat YLD2000-2d yield surface (Figure 5) comes from:
    • Conventional tensile testing at 0, 22.5, 45, 67.5, and 90 degrees to the rolling direction, determining the yield strength and the r-value;
    • Disc compression tests according to the procedure in Citation T-21 to determine the biaxial R-value, rb.

Figure 5: Tensile testing and disc compression testing generate the Barlat YLD2000-2d yield surface in two 3rd Generation AHSS Grades B-13

Figure 5: Tensile testing and disc compression testing generate the Barlat YLD2000-2d yield surface in two 3rd Generation AHSS Grades B-13


  • Creating the hardening curve uses a procedure detailed in Citations R-5 and N-13, and involves only conventional tensile and shear testing using the procedure included in Citation P-15.

Figure 6: Test geometries for hardening curve generation. Left image: Tensile; Right image: Shear.  N-13

Figure 6: Test geometries for hardening curve generation. Left image: Tensile; Right image: Shear.N-13


  • Characterizing formability involved generating a Forming Limit Curve using Marciniak data or process-corrected Nakazima data. (See our article on non-linear strain paths) and Citation N-13 for explanation of process corrections].  Either approach resulted in acceptable characterizations.
  • Fracture characterization uses four plane stress tests: shear, conical hole expansion, V-bending, and a biaxial dome test.  The result from these tests calibrate the fracture locus describing the stress states at fracture.


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Uniform Elongation

Tube Forming

Manufacturing precision welded tubes typically involves continuous roll forming followed by a longitudinal weld typically created by high frequency (HF) induction welding process known as electric resistance welding (ERW) or by laser welding.

Tubular components can be a cost-effective way to reduce vehicle mass and improve safety. Closed sections are more rigid, resulting in improved structural stiffness. Automotive applications include seat structures, cross members, side impact beams, bumpers, engine subframes, suspension arms, and twist beams. All AHSS grades can be roll formed and welded into tubes with large D/t ratios (tube diameter / wall thickness); tubes having 100:1 D/t with a 1mm wall thickness are available for Dual Phase and TRIP grades.

Figure 1: Automotive Applications for Tubular Components.A-35

Figure 1: Automotive Applications for Tubular Components.A-35



As roll formed and welded tubes are used with mounting brackets and little else in some Side Intrusion Beams (Figure 2), or they can be used as a precursor to hydroforming, such as the Engine Cradle shown in Figure 3.

Figure 2: Side intrusion beams made from a welded tube with mounting brackets.S-34

Figure 2: Side intrusion beams made from a welded tube with mounting brackets.S-34


Figure 3: The stages of a Hydroformed Engine Cradle: A) Straight Tube, B) After bending; C) After pre-forming; D) Hydroformed Engine Cradle. S-35

Figure 3: The stages of a Hydroformed Engine Cradle: A) Straight Tube, B) After bending; C) After pre-forming; D) Hydroformed Engine Cradle. S-35



The processing steps of tube manufacturing affect the mechanical properties of the tube, increasing the yield strength and tensile strength, while decreasing the total elongation. Subsequent operations like flaring, flattening, expansion, reduction, die forming, bending and hydroforming must consider the tube properties rather than the properties of the incoming flat sheet.

The work hardening, which takes place during the tube manufacturing process, increases the yield strength and makes the welded AHSS tubes appropriate as a structural material. Mechanical properties of welded AHSS tubes (Figure 4) show welded AHSS tubes provide excellent engineering properties. AHSS tubes are suitable for structures and offer competitive advantage through high-energy absorption, high strength, low weight, and cost efficient manufacturing

Figure 4: Anticipated Properties of AHSS Tubes; A) Yield Strength, B) Total Elongation.R-1

Figure 4: Anticipated Properties of AHSS Tubes; A) Yield Strength, B) Total Elongation.R-1



The degree of work hardening, and consequently the formability of the tube, depends both on the steel grade and the tube diameter/thickness ratio (D/T) as shown in Figure 5. The degree of work hardening influences the reduction in formability of tubular materials compared with the as-produced sheet material. Furthermore, computerized forming-process development utilizes the actual true stress-true strain curve of steel taken from the tube, which is influenced by the steel grade, tube diameter, and forming process.

Figure 5: Examples of true stress – true strain curves for AHSS tubes made from Dual Phase Steel with 590 MPa minimum tensile strength.A-36

Figure 5: Examples of true stress – true strain curves for AHSS tubes made from Dual Phase Steel with 590 MPa minimum tensile strength.A-36



Bending AHSS tubes follows the same laws that apply to ordinary steel tubes. Splitting, buckling, and wrinkling must all be avoided. As wall thickness and bend radius decrease, the potential for wrinkling or buckling increases.

One method to evaluate the formability of a tube is the minimum bend radius. An empirically derived formulaS-36  for the minimum Centerline Radius (CLR) considers both tube diameter (D) and total elongation (A) determined in a tensile test with a proportional test specimen, and assumes tube formation via rotary draw bending:

The formula shows that a bending radius equal to the tube diameter (1xD) requires a steel with 50% elongation. Successfully bending low elongation material needs a greater bend radius. Consider, for example, a dual phase steel grade where the elongation of a sample measured off the tube is 12.5%. Here, the minimum bend radius is 4 times the tube diameter. The tube bending method, the use and type of mandrels, and choice of lubrication may all affect the CLR.

The engineering strain on the outer surface can be estimated as the tube diameter (D) divided by twice the Centerline Radius (CLR):

For example, if you are bending a 40 mm diameter tube around a centerline radius of 100mm, the engineering strain on the outer surface is approximately 40/[2*100] = 20%. As a rough estimate, successful bending requires the tube to have a minimum elongation value from a tensile test in excess of this amount. Otherwise, a larger radius or modifications to the forming process is needed.

Springback is related to the elastic behavior of the tube. Yield strength variation between production batches can lead to variation in the amount of springback. A rule of thumb is that a variation of ± 10 MPa in yield strength causes a variation of approximately ± 0.1° in the bending radius. In addition, a high frequency weld which is harder and stronger than the base material causes a maximum of approximately ± 0.1° variation in bending radius.S-36

The bending behavior of tube depends on both the tubular material and the bending technique. The weld seam is also an area of non-uniformity in the tubular cross section, and therefore influences the forming behavior of welded tubes. The recommended procedure is to locate the weld area in a neutral position during the bending operation.

Figures 6 and 7 provide examples of the forming of AHSS tubes. The discussion on Tailored Products describes tailored tubes, which may be further hydroformed.

Figure 6: Dual phase steel bent to 45 degrees with centerline bending radius of 1.5xD using booster bending. Steel properties in the tube: 610 MPa yield strength, 680 MPa tensile strength, 27% total elongation. R-1

Figure 6: DP steel bent to 45 degrees with centerline bending radius of 1.5xD using booster bending. Steel properties in the tube: 610 MPa yield strength, 680 MPa tensile strength, 27% total elongation. R-1


Figure 7: Hydroformed Engine Cradle made from a dual phase steel welded tube by draw bending with centerline bending radius of 1.6xD and a bending angle greater than 90 degrees. Steel properties in the tube: 540 MPa yield strength, 710 MPa tensile strength, 34% total elongation. R-1

Figure 7: Hydroformed Engine Cradle made from a dual phase steel welded tube by draw bending with centerline bending radius of 1.6xD and a bending angle greater than 90 degrees. Steel properties in the tube: 540 MPa yield strength, 710 MPa tensile strength, 34% total elongation. R-1



Key Points

  • Due to the cold working generated during tube forming, the formability of the tube is reduced compared to the as-received sheet.
  • The work hardening during tube forming increases the YS and TS, thereby allowing the tube to be a structural member.
  • Successful bending requires aligning the targeted radii with the available elongation of the selected steel grade.
  • The weld seam should be located at the neutral axis of the tube, whenever possible during the bending operation.