Dr. Donald Malen, College of Engineering, University of Michigan, reviews the use of two recently developed Powertrain Models, which he co-authored with Dr. Roland Geyer, University of California, Bren School of Environmental Science.
The Need For The Powertrain Models
The use of Advanced High-Strength Steel (AHSS) grades offer a means to lightweight a vehicle. Among the benefits of this lightweighting are less fuel used over the vehicle life, and better acceleration performance. Vehicle designers as well as Greenhouse gas analysts are interested in estimating these benefits early in the vehicle design process 5.
Models are constructed for this purpose which range from the use of a simple coefficient, (for example fuel consumption change per kg of mass reduction), to very detailed models accessible only to specialists which require knowledge of hundreds of vehicle parameters. Draw backs to the first approach is that the coefficient may be based on assumptions about the vehicle which do not match the current case. Drawbacks to the detailed models are the considerable expense and time needed, and the lack of transparency in the results; It is difficult to relate inputs with outputs.
A middle way between the simplistic coefficient and the complex model, is described here as a set of Parsimonious Powertrain Models 1, 2, 3. Parsimony is the principle that the best model is the one that requires the fewest assumptions while still providing adequate estimates. These Excel spreadsheet models cover Internal Combustion powertrains, Battery Electric Vehicles, and Plug-in Electric Vehicles, and predict fuel consumption and acceleration performance based on a small set of inputs. Inputs include vehicle characteristics (mass, drag coefficient, frontal area, rolling resistance), powertrain characteristics (fuel conversion efficiency, gear ratios, gear train efficiency), and fuel consumption driving cycle. Model outputs include estimates for fuel consumption, acceleration, and a visitation map.
Physics of the Models
Fuel consumption is determined by the quantity of fuel used over a driving cycle. The driving cycle specifies the vehicle speed vs. time. An example of a driving cycle is the World Light Vehicles Test Procedure (WLTP) cycle shown in Figure 1.
Figure 1: Fuel Consumption Driving Cycle (WLTP Class 3b).
Given the velocity history of Figure 1, the forces on the vehicle resisting forward motion may be calculated. These forces include inertia force, aerodynamic drag force, and rolling resistance. The total of these forces, called tractive force, must be provided by the vehicle propulsion system, see Figure 2.
Figure 2. Tractive Force Required.
Once vehicle speed and tractive force are known at each point of time during the driving cycle, the required torque and rotational speed may be determined for each of the drivetrain elements, as shown in Figure 3 for an Internal Combustion system, and Figure 4 for a Battery Electric Vehicle.
Figure 3. Internal Combustion Powertrain.
Figure 4. Battery Electric Vehicle Powertrain.
In this way, the required torque and speed of the engine or motor may be determined. Then using a map of efficiency, shown to the right in Figures 3 and 4, the energy demand is determined at each point in time. Summing the energy demand over time yields the fuel used over the driving cycle. The reader is referred to References 1 and 2 for a much more in depth description of the models.
As an example application, consider the WorldAutoSteel FutureSteelVehicle (FSV) 4. The FSV project, completed in 2011, investigated the weight reduction potential enabled with the use of AHSS, advanced manufacturing processes and computer optimization. The resulting material use in the body structure is shown in Figure 5.
Figure 5. FutureSteelVehicle steel grade application.
This use of AHSS allowed a reduction in the vehicle curb mass from 1200 kg to 1000 kg. What are the effects of this mass reduction on fuel consumption and acceleration performance?
The inputs required for the powertrain model are shown in Table 1 for the base case.
Table 1: Model Inputs for Base Case
The results provided by the powertrain model are summarized in the acceleration-time vs. fuel consumption graph of Figure 6. Point A is the base case at 1200 kg curb mass. The lightweight case with same engine is shown as Point B. Note the fuel consumption reduction and also the acceleration time reduction. Often the acceleration time is set as a requirement. For the lighter vehicle, the engine size may be reduced to achieve the original acceleration time and an even greater reduction in fuel consumption as shown as Point C.
Figure 6. Summary of results of base vehicle and reduced mass vehicle.
Using the parsimonious powertrain models allows such ‘what-if’ questions to be answered quickly, with minimal data input, and in a transparent way. The Parsimonious Powertrain Models are available as a free download at worldautosteel.org.
1 Geyer, R., Malen, D.E. Parsimonious powertrain modeling for environmental vehicle assessments: part 1—internal combustion vehicles. Int Journal of Life Cycle Assessment (2020), JLCA-D-19-00183R1, June 22, 2020
2 Geyer, R., Malen, D.E. Parsimonious powertrain modeling for environmental vehicle assessments: part 2—electric vehicles. Int Journal of Life Cycle Assessment (2020), JLCA-D-19-00250R1, June 15, 2020,
3 Powertrain Excel Models for free download.
4 FutureSteelVehicle Results and Reports, Feb 18, 2015
5 UCSB Automotive Energy & GHG Model, Mar 14, 2017 (The Powertrain Models were used to calculate the Fuel Reduction Value for this model.)
Results of a Three-Year LME Study
WorldAutoSteel releases today the results of a three-year study on Liquid Metal Embrittlement (LME), a type of cracking that is reported to occur in the welding of Advanced High-Strength Steels (AHSS).The study results add important knowledge and data to understanding the mechanisms behind LME and thereby finding methods to control and establish parameters for preventing its occurrence. As well, the study investigated possible consequences of residual LME on part performance, as well as non-destructive methods for detecting and characterizing LME cracking, both in the laboratory and on the manufacturing line (Figure 1).
Figure 1: LME Study Scope
The study encompassed three different research fields, with an expert institute engaged for each:
- Laboratory of Materials and Joining Technology (LWF) Paderborn University, Paderborn, Germany, for experimental research,
- The Institute de Soudure, Yutz, France, for investigations regarding non-destructive testing, and
- Fraunhofer Institute for Production Systems and Design Technology (IPK), Berlin, Germany, for the development of a simulation model to accurately replicate welding conditions resulting in LME.
A portfolio containing 13 anonymized AHSS grades, including dual phase (DP), martensitic (MS) and retained austenite (RA) with an ultimate tensile strength (UTS) of 800 MPa and higher, was used to set up a testing matrix, which enabled the replication of the most relevant and critical material thickness combinations (MTC). All considered MTCs show a sufficient weldability under use of standard parameters according to SEP1220-2. Additional MTCs included the joining of various strengths and thicknesses of mild steels to select AHSS in the portfolio. Figure 2 provides the welding parameters used throughout the study.
Figure 2: LME Study Welding Parameters
In parallel, a 3D electro-thermomechanical simulation model was set up to study LME. The model is based on temperature-dependent material data for dual phase AHSS as well as electrical and thermal contact resistance measurements and calculates local heating due to current flow as well as mechanical stresses and strains. It proved particularly useful in providing additional means to mathematically study the dynamics observed in the experimental tests. This model development was documented in two previous AHSS Insights blogs (see AHSS Insights Related Articles below).
The study began by analyzing different influence factors (Figure 3) which resembled typical process deviations that might occur during car body production. The impact of the influences was analyzed by the degree of cracking observed for each factor. A select number of welding set-ups from these investigations were rebuilt digitally in the simulation model to replicate the process and study its dynamics mathematically. This further enabled the clarification of important cause-effect relationships.
Figure 3: Overview of All Applied Influence Factors (those outlined in yellow resulted in most frequent cracking.)
Generally, the most frequent cracking was observed for sharp electrode geometries, increased weld times and application of external loads during welding. All three factors were closely analyzed by combining the experimental approach with the numerical approach using the simulation model.
Destructive Testing – LME Effects on Mechanical Joint Strength
A destructive testing program also was conducted for an evaluation of LME impact on mechanical joint strength and load bearing capacity in multiple conditions, including quasi-static loading, cyclic loading, crash tests and corrosion. In summary of all load cases, it can be concluded that LME cracks, which might be caused by typical process deviations (e.g. bad part fit up, worn electrodes) have a low intensity impact and do not affect the mechanical strength of the spot weld. And as previously mentioned, the study analyses showed that a complete avoidance of LME during resistance spot welding is possible by the application of measures for reducing the critical conditions from local strains and exposure to liquid zinc.
In welding under external load experiments, the locations of the experimental crack occurrence showed close correlation with the strains and remaining plastic deformations computed by the simulation model. It was observed that the cracks form at the location of the highest plastic strains, and material-specific threshold values for critical strains were derived. The threshold values then were used to judge the crack formation at elongated weld times.
At the same time, the simulation model pointed out a significant difference in liquid zinc diffusion during elongated weld times. Therefore, it is concluded that liquid zinc exposure time is a second highly relevant factor for LME formation.
The results for the remaining influence factors depended on the investigated MTCs and were generally less significant. In more susceptible MTCs (AHSS welded with thick Mild steel), no significant cracking occurred when welded using standard process parameters. Light cracking was observed for most of the investigated influences, such as low electrode cooling rate, worn electrode caps, electrode positioning deviations or for gap afflicted spot welds. More intense cracking (higher penetration depth cracking) was only observed when welding under extremely high external loads (0.8 Re) or, even more, as a consequence of highly increased weld times.
For the non-susceptible MTCs, even extreme situations and weld set-ups (such as the described elongated weld times) did not result in significant LME cracks within the investigated AHSS grades.
Methods for avoidance of LME also were investigated. Changing the electrode tip geometry to larger working plane diameters and elongating the hold time proved to eliminate LME cracks. In the experiments, a change of electrode tip geometry from a 5.5 mm to an 8.0 mm (Figure 4) enabled LME-free welds even when doubling the weld times above 600 ms. Using a flat-headed cap (with small edge radii or beveled), even the most extreme welding schedules (weld times greater than 1000 ms) did not produce cracks. The in-depth analysis revealed that larger electrode tip geometries clearly reduce the local plastic deformation around the indentation. This plastic strain reduction is particularly important, as longer weld times contribute to a higher liquid zinc exposure interval, leading to a higher potential for LME cracks.
Figure 4: Electrode Geometries Used in Study Experiments
It was also seen that as more energy flows into a spot weld, it becomes more critical to parameterize an appropriate hold time. Depending on the scenario, the selection of the correct hold time alone can make the difference between cracked and crack-free welds. Insufficient hold times allow liquid zinc to remain on the steel surface and increased thermal stresses that form after the lift-off of the electrode caps. Elongated hold times reduce surface temperatures, minimizing surface stresses and thus LME potential.
NON-Destructive Testing: Laboratory and Production Capabilities
A third element of the study, and an aid in the control of LME, is the detection and characterization of LME cracks in resistance spot welds, either in laboratory or in production conditions. This work was done by the Institute of Soudure in close cooperation with LWF, IPK and WorldAutoSteel members’ and other manufacturing facilities. Ten different non-destructive techniques and systems were investigated. These techniques can be complementary, with various levels of costs, with some solutions more technically mature than others. Several techniques proved to be successful in crack detection. In order to aid the production source, techniques must not only detect but also characterize cracks to determine intensity and the effect on joint strength. Further work is required to achieve production-level characterization.
The study report provides detailed technical information concerning the experimental findings and performances of each technique/system and the possible application cost of each. Table 1 shows a summary of results:
Table 1: Summary of NDT: LME Detection and Characterization Methods
Suitable measures should always be adapted to the specific use case. Generally, the most effective measures for LME prevention or mitigation are:
- Avoidance of excessive heat input (e.g. excess welding time, current).
- Avoidance of sharp edges on spot welding electrodes; instead use electrodes with larger working plane diameter, while not increasing nugget-size.
- Employing extended hold times to allow for sufficient heat dissipation and lower surface temperatures.
- Avoidance of improper welding equipment (e.g. misalignments of the welding gun, highly worn electrodes, insufficient electrode cooling)
In conclusion, a key finding of this study is that LME cracks only occurred in the study experiments when there were deviations from proper welding parameters and set-up. Ensuring these preventive measures are diligently adhered to will greatly reduce or eliminate LME from the manufacturing line. For an in-depth review of the study and its findings, you can download a copy of the full report at worldautosteel.org.
LME Study Authors
The LME study authors were supported by a committed team of WorldAutoSteel member companies’ Joining experts, who provided valuable guidance and feedback.
Related Articles: More on this study in previous AHSS Insights blogs:
- Julian Frei, Max Biegler, Michael Rethmeier, Christoph Böhne & Gerson Meschut (2019) Investigation of liquid metal embrittlement of dual phase steel joints by electro-thermomechanical spot-welding simulation, Science and Technology of Welding and Joining, 24:7, 624-633, DOI: 10.1080/13621718.2019.1582203
- Christoph Böhne, Gerson Meschut, Max Biegler, Julian Frei & Michael Rethmeier (2020) Prevention of liquid metal embrittlement cracks in resistance spot welds by adaption of electrode geometry, Science and Technology of Welding and Joining, 25:4, 303-310, DOI: 10.1080/13621718.2019.1693731
In static or dynamic conditions, the spot weld strength of Advanced High-Strength Steels (AHSS) may be considered as a limiting factor. One solution to improve resistance spot weld strength is to add a high-strength adhesive to the weld. Figure 1 illustrates the strength improvement obtained in static conditions when crash adhesive (in this case, Betamate 1496 from Dow Automotive) is added. The trials were performed with 45-mm-wide and 16-mm adhesive bead samples.
Figure 1: Tensile Shear Strength and Cross Tensile Strength on DP 600.1
Another approach to improve the strength of welds is done by using laser welding instead of spot welding. Compared to spot welding, the main advantage of laser welding, with respect to the mechanical properties of the joint, is the possibility to adjust the weld dimension to the requirement. One may assume that, in tensile shear conditions, the weld strength depends linearly on the weld length as indicated in the results of a trial 1, shown in Figure 2.
Figure 2: Tensile-shear strength on laser weld stitches of different length.1
However, a comparison of spot weld to laser weld strength cannot be restricted to the basic tensile shear test. Tests were also conducted to evaluate the weld strength in both quasi-static and dynamic conditions under different solicitations, on various AHSS combinations. The trials were performed on a high-speed testing machine, at 5 mm/min for the quasi-static tests and 0.5 m/s for the dynamic tests (pure shear, pure tear or mixed solicitation, as shown in Figure 3). The strength at failure and the energy absorbed during the trial were measured. Laser stitches were done at 27mm length. C- and S-shape welds were performed with the same overall weld length.
Figure 3: Sample geometry for quasi-static and dynamic tests.1
The weld strength at failure is described in Figure 4, where major axes represent pure shear and tear (Figure 4). For a reference spot weld corresponding to the upper limit of the weldability range, globally similar weld properties can be obtained with 27mm laser welds. The spot weld equivalent length of 25-30 mm has been confirmed on other test cases on AHSS in the 1.5- to 2 mm thickness range. It has also been noticed that the spot weld equivalent length is shorter on thin mild steel (approximately 15-20 mm). This must be considered when shifting from spot to laser welding on a given structure. There is no major strain rate influence on the weld strength; the same order of magnitude is obtained in quasi-static and dynamic conditions.
Figure 4: Quasi-static and dynamic strength of welds, DP 600 2 mm+1.5 mm.1
The results in terms of energy absorbed by the sample are seen in Figure 5. In tearing conditions, both the strength at fracture and energy are lower for the spot weld than for the various laser welding procedures. In shear conditions, the strength at fracture is equivalent for all the welding processes. However, the energy absorption is more favorable to spot welds. This is due to the different fracture modes of the welds; for example, interfacial fracture is observed on the laser welds under shearing solicitation. Even if the strength at failure is as high as for the spot welds, this severe failure mode leads to lower total energy absorption.
Figure 5: Strength at fracture and energy absorption of Hot Rolled 1500 1.8-mm + DP 600 1.5-mm samples for various welding conditions.1
Figure 6 represents the energy absorbed by omega-shaped structures and the corresponding number of welds that fail during the frontal crash test (here on TRIP 800 grade). It appears clearly that laser stitches have the highest rate of fracture during the crash test (33%). In standard spot welding, some weld fractures also occur. It is known that AHSS are more prone to partial interfacial fracture on coupons, and some welds fail as well during crash tests. By using either Weld-Bonding or adapted laser welding shapes, weld fractures are mitigated, even in the case of severe deformation. As a consequence, higher energy absorption is also observed.
Figure 6: Welding process and weld shape influence on the energy absorption
and weld integrity on frontal crash tests. 1
Up to a 20% improvement can be achieved in torsional stiffness, where the best results reflected the combination of laser welds and adhesives. Adhesive bonding and weld- bonding lead to the same stiffness improvement results due to the adhesive rather than the additional welds. Figure 7 shows the evolution of the torsional stiffness with the joining process. Optimized laser joining design leads to the same performances as a weld bonded sample in fracture modes, shown in Figure 8.
Figure 7: Evolution of the torsional stiffness with the joining process.1
Figure 8: Validation test case 1.2-mmTRIP 800/1.2-mm hat-shaped TRIP 800.
Top-hat crash boxes were tested across a range of AHSS materials including DP 1000. The spot weld’s energy absorption increased linearly with increasing material strength. The adhesives were not suitable for crash applications as the adhesive peels open along the entire length of the joint. The weld bonded samples perform much better than conventional spot welds. Across the entire range of materials there was a 20-30% increase in mean force when weld bonding was used; the implications suggesting a similarly significant improvement in crash performance. Furthermore, results show that a 600 MPa weld bonded steel can achieve the same crash performance as a 1000 MPa spot-welded steel. It is also possible that some down gauging of materials could be achieved, but as the strength of the crash structure is highly dependent upon sheet thickness, only small gauge reductions would be possible. Figure 9 shows the crash results for spot-welded and weld bonded AHSS.
Figure 9: Crash results for spot-welded and weld bonded AHSS.
1 Courtesy of ArcelorMittal.
Many steel parts on a vehicle require corrosion protection, regardless of whether they are exposed or unexposed applications. The most common way to accomplish corrosion protection is to coat Advanced High-Strength Steels (AHSS) with zinc by means of a couple of different processes. This AHSS Insights Blog goes over the most common.
Electrogalvanizing is a zinc deposition process, where the zinc is electrolytically bonded to steel in order to protect against corrosion. The process involves electroplating: running an electrical current through the steel strip as it passes through a saline/zinc solution. Electrogalvanizing is done at room temperature, so the microstructure, mechanical, and physical properties of AHSS products achieved on a continuous anneal line (CAL) are essentially unchanged after the electrogalvanizing (EG) process. EG lines have multiple plating cells, with each cell capable of being on or off. As a result, chief advantages of electrogalvanizing compared to hot dipped galvanizing include: (1) lower processing temperatures, (2) precise coating weight control, and (3) brighter, more uniform coatings which are easier to convert to Class A exposed quality painted surfaces.
The majority of electrogalvanizing lines can apply only pure (free) zinc coatings, known as EG for electrogalvanized steel. Selected lines can apply different types of coatings, like EGA (electro-galvanneal) or Zn-Ni (zinc-nickel).
There are no concerns about different alloy phases in the coating as with galvanneal coatings. The lack of aluminum in the coating results in improved weldability. The biggest concern with electrogalvanizing lines is the coefficient of friction. Electrogalvanized (EG) coatings have a relatively high coefficient of friction—higher than hot dipped galvanized coatings, but lower than galvanneal coatings. To improve formability of electrogalvanized sheets, some automakers choose to use a steel mill-applied pre-lube rather than a simple mill-applied rust preventive oil.
A representative EG line is shown in Figure 1. Different EG lines may use different technologies to apply the zinc crystals. Because the zinc crystals are deposited in a different fashion, these different processes may potentially result in different surface morphology and, in turn, a different coefficient of friction.
Figure 1: Schematic of an electrogalvanizing line.
A higher coefficient of friction may be found under dry conditions, but the “stacked plate-like surface morphology” (Figure 2) allows these coatings to trap and hold lubrication better than the smoother surfaces of hot dipped galvanizing coatings. Auto manufacturers should therefore consult the steel supplier for specific lubricant recommendations based on the forming needs.
Figure 2: High magnification photograph of electrogalvanized steel surface showing stacked plate-like structure.
Hot Dip Galvanize and Hot Dip Galvanneal
Hot dipped galvanizing – applying a zinc coating over the steel – is the most common way to achieve corrosion protection. It is an economical solution, since cold rolled steel can be annealed and coated in the same continuous operation.
A typical in-line continuous hot dip galvanizing line such as that shown in Figure 3 uses a full-hard cold rolled steel coil as the feedstock. Individual coils are welded together to produce a continuous strip. After cleaning, the strip is processed in a continuous annealing furnace where the microstructure is recrystallized, improving forming characteristics. The annealing temperature is adjusted to produce the desired microstructure associated with the ordered grade. Rather than cooling to room temperature, the in-process coil is cooled to just above 460°C (860°F), the temperature of the molten zinc bath it enters. The chemistry in the zinc pot is a function of whether a hot dipped galvanized or galvannealed coating is ordered. Hot rolled steels also are coated with the hot dip galvanizing process, but different processing conditions are used to achieve the targeted properties.
Figure 3: Schematic of a typical hot dipped galvanizing line with galvanneal capability.
There are several types of hot dipped coatings for automotive applications, with unique characteristics that affect their corrosion protection, lubricity for forming, weldability and paintability. One of the primary hot dipped galvanized coatings is a pure zinc coating (abbreviated as GI), sometime referred to as free zinc. The molten zinc bath has small amounts of aluminum which helps to form a thin Fe2Al5 layer at the zinc-steel interface. This thin barrier layer prevents zinc from diffusing into the base steel, which leaves the coating as essentially pure zinc.
Coil pass through the molten zinc at speeds up to 3 meters per second. Zinc coating weight is controlled by gas knives (typically air or nitrogen) blowing off excess liquid zinc as the coil emerges from the bath. Zinc remaining on the surface solidifies into crystals called spangle. Molten zinc chemistry and cooling practices used at the galvanizing line control spangle size. Since spangle can show through on a painted surface, a minimum-spangle or no-spangle option is appropriate for surface-critical applications.
The other primary hot dipped coating used for corrosion protection is hot dipped galvanneal (abbreviated as GA). Applying this coating to a steel coil involves the same steps as creating a free zinc hot dipped coated steel, but after exiting the zinc pot, the steel strip passes through a galvannealing furnace where the zinc coating is reheated while still molten.
The molten zinc bath used to produce a GA coating has a lower aluminum content than what is used to produce a GI coating. Without aluminum to create the barrier layer, the zinc coating and the base steel inter-diffuse freely, creating an iron-zinc alloy with typical average iron content in the 8-12% range. The iron content improves weldability, which is a key attribute of the galvanneal coatings.
The iron content will be unevenly distributed throughout the coating, ranging from 5% at the surface (where the sheet metal coating contacts the tool surface during forming) to as much as 25% iron content at the steel/coating interface. The amount of iron at the surface and distribution within the coating is a function of galvannealing parameters and practices – primarily the bath composition and time spent at the galvannealing temperature. Coating iron content impacts coating hardness, which affects the interaction with the sheet forming lubricant and tools, and results in changes in friction. The hard GA coatings have a greater powdering tendency during contact with tooling surfaces, especially during movement through draw beads. Powdering is minimized by using thinner coatings – where 50 g/m2 to 60 g/m2 (50G to 60G) is a typical EG and GI coating weight, GA coatings are more commonly between 30 g/m2 to 45 g/m2 (30A to 45A).
Figure 4: High magnification photograph of a galvannealed steel surface. The surface structure results in excellent paint adhesion.
Options to improve formability on parts made from GA coated steels include use of press-applied lubricants or products that can be applied at the steel mill after galvanizing, like roll-coated phosphate, which have the additional benefit of added lubricity. The surface morphology of a galvannealed surface (Figure 4) promotes good phosphate adherence, which in turn is favorable for paintability.
Galvannealed coatings provide excellent corrosion protection to the underlying steel, as do GI and EG coatings. GI and EG coatings are essentially pure zinc. Zinc acts as a sacrificial anode if either coating is damaged from scratches or impact, and therefore will corrode first before the underlying steel. The corrosion product of GI and EG is white and is a combination of zinc carbonate and zinc hydroxide. A similar mechanism protects GA coated steels, but the presence of iron in the coating may result in a reddish tinge to the corrosion product. This should not be interpreted as an indication of corrosion of the steel substrate.
Producing galvanized and galvannealed AHSS is challenging due to the interactions of the necessary thermal cycles at each step. As an example, the targeted microstructure of Dual Phase steels can be achieved by varying the temperature and time the steel strip passes through the zinc bath and can be adjusted to achieve the targeted strength level. However, not all AHSS can attain their microstructure with the thermal profile of a conventional hot dipped galvanizing line with limited rapid quenching capabilities. In addition, many AHSS grades have chemistries that lead to increased surface oxides, preventing good zinc adhesion to the surface. These grades must be produced on a stand-alone Continuous Annealing Line, or CAL, without an in-line zinc pot. Continuous Annealing Lines feature a furnace with variable and rapid quenching operations that enable the thermal processing required to achieve very high-strength levels. If corrosion protection is required, these steel grades are coated on an electrogalvanizing line (EG) in a separate operation, after being processed on a CAL line.
Hot dipped galvanizing lines at different steel companies have similar processes that result in similar surfaces with respect to coefficient of friction. Surface finish and texture (and resultant frictional characteristics) are primarily due to work roll textures, based on the customer specification. Converting from one coating line to another using the same specification is usually not of major significance with respect to coefficient of friction. A more significant change in friction is observed with changes between GI and GA and EG.
In the 2nd Quarter 2020, we’ll release the results of a three-year study on Liquid Metal Embrittlement in resistance spot welding. It will shed light on why LME occurs, how its occurrence can be controlled, and practical preventative measures to avoid LME on the manufacturing line. Stay tuned.
In this edition of AHSS Insights, George Coates and Menachem Kimchi get back to basics with important fundamentals in forming and joining AHSS.
As the global steel industry continues its development of Advanced High-Strength Steels (AHSS), including 3rd Gen products with enhanced formability, we’re reminded that successful application is still dependent on the fundamentals, both in forming and joining. In this blog article, we address some of those forming considerations, as well as highlighting common joining issues in manufacturing.
The somewhat lower formability of AHSS compared to mild steels can almost always be compensated for by modifying the design of the component and optimizing blank shape and the forming process.
In stamping plants, we’ve observed inconsistent practices in die set-up and maintenance, surface treatments and lubrication application. Some of these inconsistencies can be addressed through education, via training programs on AHSS Application Guidelines. These Guidelines share best practices and lessons learned to inform new users on different behaviors of specific AHSS products, and the necessary modifications to assist their application success. In addition to new practices, we’ve learned that applying process control fundamentals become more critical as one transitions from mild steels to AHSS, because the forming windows are smaller and less forgiving, meaning these processes don’t tolerate variation well. If your present die shop is reflective of housekeeping issues, such as oil and die scrap on the floor or die beds, you are a candidate for a shop floor renovation or you will struggle forming AHSS products.
Each stamping operation combines three main elements to achieve a part meeting its desired functional requirements:
- the steel product,
- appropriate die materials, including their surface treatment, and
- the correct lubricant that maintains its lubricity during the forming operation.
There is good news, in that our industry is responding with new products and services to improve manufacturing performance and save costs.
As an example, lubrication application is often overlooked, and old systems may be ineffective. In the forming of AHSS, part temperatures can become excessive, and break down lubricant performance. Figure 1 shows an example of part temperatures from an Ohio State University study conducted with DP 980 steels1.
Figure 1: Example Temperature distribution for DP 980 Steel1.
Stampers often respond by “flooding” the process with extra lubricant, thinking this will solve their problem. Instead, lubricant viscosity and high temperature stability are the most important considerations in the lubricant selection, and new types exist to meet these challenges. Also, today there are new lubrication controllers that can finely control and disperse wet lubricants evenly across the steel strip, or in very specific locations, if forming requirements are localized. These enable better performance while minimizing lubricant waste (saving cost), a win-win for the pressroom.
Similarly, AHSS places higher demands on tool steels used in forming and cutting operations. In forming applications, galling, adhesive wear and plastic deformation are the most common failure mechanisms. Surface treatments such as PVD, CVD and TD coatings applied to the forming tool are effective at preventing galling. Selection of the tool steel and coating process used for forming AHSS will largely depend on the:
- Strength and thickness of the AHSS product,
- Steel coating,
- Complexity of the forming process, and
- Number of parts to be produced.
New die materials such as “enhanced D2” are available from many suppliers. These improve the balance between toughness, hardness and wear resistance for longer life. These materials can be thru-hardened, and thus become an excellent base material for PVD or secondary surface treatments necessary in the AHSS processing. And new tool steels have been developed specifically for hot forming applications.
In high-volume production different Resistance Spot Welding (RSW) process parameters can be used depending on the application and the specifications applied. Assuming you chose the appropriate welding parameters that allows for a large process window, manufacturing variables may ruin your operation as they strongly effect the RSW weld quality and performance.
One of the great advantages of the RSW process is the action of clamping the material together via the electrode force applied during the process. However due to the pre-welding condition/processing such as the stamping operation, this fit-up issue, as shown in Figure 2, can be very significant especially in welding an AHSS product. In this case the effective required force specified during the process setup for the application is significantly reduced and can result in an unacceptable weld, over-heating, and severe metal expulsion. If the steels are coated, higher probability for Liquid Metal Embrittlement (LME) cracking is possible.
Figure 2: Examples of Pre-Welding Condition/Processing Fit-Up Issues
For welding AHSS, higher forces are generally required as a large part of the force is being used to force the parts together in addition to the force required for welding. Also, welding parameters may be set for pre-heating with lower current pulses or current up-slope to soften the material for easier material forming and to close the gap.
During machine set up, the RSW electrodes need to be carefully aligned as shown in Figure 3A. However, in many production applications, electrode misalignment is a common problem.
Electrode misalignment in the configurations shown in Figure 3B may be detrimental to weld quality of any RSW application. Of course, the electrode misalignment shown in this figure is exaggerated but the point is that it happens frequently on manufacturing welding lines.
Figure 3: Alignment vs. Misalignment of Electrodes
In these cases, the intendent contact between the electrodes is not achieved and thus the current density and the force density (pressure) required for producing an acceptable weld cannot be achieved. With such conditions, overheating, expulsion, sub-size welds and extensive electrode wear will result. Again, if coated steels are involved, the probability for LME cracking is higher.
Note also that following specifications or recommendations for water cooling the electrode is always important for process stability and extending electrode life.
Figure 4: Sequence of Squeeze Time and Welding Current Initiation
The squeeze time is the time required for the force to reach the level needed for the specific application. Welding current should be applied only after reaching this force, as indicated in Figure 4. All RSW controllers enable the easy control of squeeze time, just as with the weld time, for example. In many production operations, a squeeze time is used that is too low due to lack of understanding of its function. If squeeze time is too low, high variability in weld quality in addition to severe expulsion will be the result.
The squeeze time required for an application depends on the machine type and characteristics (not an actual welding parameter such as weld time or welding current for example).
Some of the more modern force gauges have the capability to produce the curve shown in the Figure so adequate squeeze time will be used. If you do not know what the required squeeze time for your machine/application is, it is recommended to use a longer time.
For more on these topics, download the free AHSS Application Guidelines and/or browse the related blog topics in the menu at the right.
Source: 1 Courtesy of Ohio State University
Technical Director, WorldAutoSteel and
The Phoenix Group
Since 1991, George has been providing engineering and consulting services for industry leaders in the steel, automotive, and manufacturing industries. George’s areas of expertise include: management and strategic consulting, project management, metal processing and stamping throughput improvement, metal formability and reference panel systems, and new vehicle launch manufacturing support. George is an active contributor to WorldAutoSteel technical programs, including project director / instructor for AHSS Application Guidelines.
Associate Professor, Dept of Materials Science, Ohio State University
Technical Editor, AHSS Application Guidelines
Menachem Kimchi is a associate professor in the Department of Material Science & Engineering, Welding Engineering Program at The Ohio State University. Previously, M. Kimchi was a principal research engineer, technology leader, and business development manager for EWI Manufacturing. Menachem has been involved extensively with development projects for the Automotive and Steel industries and published over 100 technical papers in area of Resistance and Solid State Welding processes of advanced materials. He currently serves as the Joining Technical Editor on the AHSS Application Guidelines.
Happy New Year! We are pleased to provide this contribution by Dr. Daniel Schaeffler, President, Engineering Quality Solutions, Inc. and Technical Editor, AHSS Applications Guidelines.
Forty years ago, the metal forming community needed to figure out how to stamp a new exotic family of steels making inroads into automotive body construction. These grades, called High Strength Low Alloy steels, were much stronger than the commonplace mild steels, and were more formable than the high-strength options available at that time. Initially, only a few steelmakers were able to offer these new grades, but over time more companies added the equipment and know-how necessary to support their customers with these products. Automakers and their supply chain stampers needed to adapt as an increasing number of parts transitioned to HSLA steels.
Fast-forward a few decades, and metal formers are facing similar challenges. Successful forming and joining of Advanced High-Strength Steels is made easier with processes that are tuned to work with the characteristics associated with these alloys. One such technique to improve formability is to employ Active Binder Force Control.
In conventional stamping, a draw ring applies pressure around the binder in order to control the sheet metal flow into the cavity. The ring may be referred to as a binder plate, draw pad, pressure pad, or blank holder. Creating the restraining force typically is done with urethane springs, coil springs, gas springs (like air or nitrogen), or press cushion systems actuated by gas or hydraulic cylinders.
Where the traditional approach applies binder pressure uniformly throughout the press stroke, modern stamping presses can be equipped with cushions having multipoint-control systems (see Figure 1 example). The associated pressure profile can be adjusted around the panel and throughout the stroke to optimize metal flow, prevent splits and wrinkles, and minimize thinning.
Figure 1. An Example of Multi-Point Press Forming Method1
Incorporating Active Binder Control capabilities has several benefits for the press shop, panel quality, and product design, including:
- A segmented blankholder combined with individually programmable hydraulic cylinders, sometimes called a flexible binder, allows for precise control of one segment independent of the others.
- Pulsating blank holder force has been shown to reduce press tonnage requirements and increase metal flow, with the frequency and amplitude being key variables that must be adjusted based on the grade and thickness of interest.
- Pre-acceleration of the cushion reduces shock loading, which minimizes the press-damaging snap-through loads associated with reverse tonnage.
The merits of a variable blank holder force on AHSS springback were documented in a 2004 conference paper2. With the traditional constant binder force approach, springback in the form of side-wall curl was seen in parts made from either a DP590 grade or a mild steel grade used as a control. Increasing the constant binder force helped to reduce springback in the mild steel part.
In Figure 2, CBF reflects tests conducted with constant binder force. VBF-LH and VBF-MLH reflect variable binder force tests conducted with a low-high force profile sequence and a medium-low-high force profile, respectively.
Figure 2: Variable Binder Force Reduces Springback2
By employing a variable binder force, springback of both the mild steel and the DP 590 material was substantially reduced. Employing either variable binder force approach reduced the thinning from forming the DP 590 material, resulting in a more uniform strain distribution across the entire channel profile (Figure 3).
Figure 3: Uniform Strain Distribution Achieved with Variable Binder Force2
More recently, a presentation from 2018 showed CP 800 panel quality improvements associated with variable blank holder force capabilities3. Panel results from a constant binder force of 300 kN and 400 kN are shown in Figures 4 and 5, respectively. Both exhibit severe wrinkling in the flange. Applying 500 kN binder force was not feasible due to exceeding the press tonnage curve limits throughout the stroke.
Figure 4: CP800 Panel Formed with Constant Binder Force of 300 kN, and Associated Close-Up of Flange3
Figure 5: CP800 Panel Formed with Constant Binder Force of 400 kN, and Associated Close-Up of Flange3
Figure 6 shows the panel produced with a variable binder force. The chosen profile fit within the press tonnage requirements and minimized wrinkles.
Figure 6: CP800 Panel Formed with Variable Binder Force Ramping from 300 kN to 600 kN, and Associated Close-Up of Flange3
Active drawbead control is an offshoot of these techniques, allowing for the magnitude and timing of drawbead engagement to be optimized for the requirements of each part. A description of using stake beads to minimize springback is available in a previous AHSS Insights blog here – active drawbead control is one approach to actuate beads.
The initial laboratory studies relating to active binder force control go back nearly 20 years ago. In the coming years, more information will enter the public domain on how metal formers are using these concepts in production. When you look to purchase a servo press, be sure to ask your press manufacturer about programmable cushions.
Daniel J. Schaeffler, Ph.D., President, Engineering Quality Solutions, Inc.;
Chief Content Officer, 4M Partners, LLC
Danny Schaeffler is the Metallurgy and Forming Technical Editor of the next release of the AHSS Applications Guidelines available from WorldAutoSteel. Danny is co-founder of 4M Partners, LLC and founder and President of Engineering Quality Solutions (EQS). He has written for Stamping Journal and The Fabricator. Danny writes the monthly “Science of Forming” column for Metalforming Magazine. He is passionate about training the metal stamping community on sheet formability and how to make the best use out of their chosen sheet metal grades.
1 Flexible Tooling For Manufacturing 3d Panels Using Multi-Point Forming Methodology – Scientific Figure on ResearchGate.
2 Variable Binder Force for Springback Management, M. Milititsky (University of Ghent), M. Garnett, C. Du, J. Wu, L. Zhang, P.J. Belanger, J.M. Prencipe (DaimlerChrysler Corporation), and E.D. Bishop (Engineering Quality Solutions, Inc.). International Conference on Advanced High-Strength Sheet Steels for Automotive Applications Proceedings, June 6-9, 2004, Association for Iron & Steel Technology
3 Improving the Drawing Process of AHSS by Using Servo Press Technologies, David Diaz-Infante (Ohio State University), 2018 Great Designs in Steel Seminar, AISI