What is Deep Drawing?

 

Deep Drawing

Introduction:

A process in which a punch forces a flat sheet metal blank into a die cavity is called a deep drawing. Through this process we can also produce shallow or moderate depth components parts. It is important metal working process because it is widely used in manufacturing.  Normally, this process is useful in producing deep parts with relatively simple shapes, with high production rate. There are many parts made by this technique like pots, pans, all the types of containers for food and beverages. The tooling and equipment cost is relatively high.           

This sheet metal forming process designed to produce hollow shells, was developed in the mid-19th century. The company started producing porcelain-enamel-covered iron cookware between 1863 and 1870, one of the first U.S. companies to do so.

Operations:

Drawing of a cup-shaped part is the basic drawing operation, with dimensions and parameters as pictured in Figure 1. A blank of diameter D_b is drawn into a die cavity by means of a punch with diameter D_p. The punch and die must have corner radii, given by R_p and R_d. If the punch and die were to have sharp corners (R_p and R_d = 0), a hole-punching operation (and not a very good one) would be accomplished rather than a drawing operation. The sides of the punch and die are separated by a clearance c. This clearance in drawing is about 10% greater than the stock thickness:

C = 1.1*t

The punch applies a downward force F to accomplish the deformation of the metal, and a downward holding force F_h is applied by the blank holder, as shown in the sketch.

As the punch proceeds downward toward its final bottom position, the work experiences a complex sequence of stresses and strains as it is gradually formed into the shape defined by the punch and die cavity. The stages in the deformation process are illustrated in Figure 2. As the punch first begins to push into the work, the metal is subjected to a bending operation. The sheet is simply bent over the corner of the punch and the corner of the die, as in Figure 2[2].

FIGURE 1 (a) Drawing of a cup shaped part: (1) start of operation before punch contacts work, and (2) near end of stroke; and (b) corresponding work part: (1) starting blank, and (2) drawn part. Symbols: C = clearance, D_b = blank diameter, D_p = punch diameter, R_d = die corner radius, R_p = punch corner radius, F = drawing force, F_h = holding force.

Deep Drawability:

In a deep drawing operation, failure generally results from the thinning of the cup wall under high longitudinal tensile stresses. If we follow the movement of the material as it flows into the die cavity, it can be seen that the sheet metal

(a)    Must be capable of undergoing a reduction in width due to a reduction in diameter.

(b)   Must also resist thinning under the longitudinal tensile stresses in the cup wall.

It has been observed that materials with outstanding deep drawability behave anisotropically. Plastic deformation in the surface is much more pronounced than in the thickness. The lankford coefficient (r) is a specific material property indicating the ratio between width deformation and thickness deformation in the uniaxial tensile test. Materials with very good deep drawability have an r value of 2 or below.

Limiting Drawing Ratio (LDR):

Deep drawability generally is expressed by the limiting drawing ratio (LDR) and the drawability of a metal is measured by the ratio of the initial blank diameter to the diameter of the cup drawn from the blank (usually approximated by the punch diameter).

 The theoretical upper limit on LDR is

Where  is an efficiency term to account for frictional losses. If  = 1, then LDR = 2.7, while if  = 0.7, LDR = 2. This agrees with experience that even with ductile metals it is difficult to draw a cup with a height much greater than its diameter [1].

So for a given material the limiting draw ratio (LDR), represents the largest blank that can be drawn through a die D_p without tearing.

Some of the practical considerations which affect drawability are[1]:

·         Die radius - should be about 10 times sheet thickness.

·         Punch radius - sharp radius leads to local thinning and tearing.

·         Clearance between punch and die ~20 to 40 percent greater than the sheet thickness.

·         Hold-down pressures about 2 per cent of average of S_o and S_u.

·         Lubricate die side to reduce friction in drawing.

In order to have successful drawing of cup shaped part has been found to be a function of the normal anisotropy, R (also called plastic anisotropy), of the sheet metal.

The effect of planar anisotropy on LDR:

Due to rolling of sheet, grains elongate into specific directions as shown in Fig. 3, so show different mechanical properties in different directions, which is called anisotropy.

Figure 3. Rolling produce smaller and elongated grains.

The planar anisotropy of the sheet is indicated by R. It is defined in terms of directional R,

The planar anisotropy decrease the LDR value, which in result cause earing defect. In deep drawing, earing defect is the formation of wavy cup of edges as shown in Fig. 4. They are objectionable on deep drawn cups because they have to be trimmed off, as they serve no useful purpose and interfere with further processing of cup, resulting in scrap.

Figure 3. Earing produced due to planar anisotropy.

The value of the strain hardening exponent (n) lies between 0 and 1. Most metals have an n value between 0.10 and 0.50. What is this value representing?

The response of a metal to cold working can be quantified by the strain-hardening exponent n. The relationship between true stress 𝞂, true strain ε, and the strain-hardening exponent n is governed by so-called power law behavior according to

𝞂 = K*εⁿ

The constant K (known as the strength coefficient) is equal to the stress when ε = 1. The value of the strain hardening exponent (n) lies between 0 and 1. Most metals have an n value between 0.10 and 0.50 [3].

It is actually a measure of the ability of a metal to strain harden; the larger its magnitude, the greater is the strain hardening for a given amount of plastic strain [4].  The value of 0 means there is no strain hardening which indicates that it is perfectly plastic solid, greater than 0 means a little strain hardening which most of the metals shows, while value of 1 indicates that true stress and true strain are linearly dependent which indicates that it is 100% elastic solid.

The effect of grain-size of sheet metal on the surface finish of the cup:

Grain size of sheet depend on many factors, rolling temperature, composition and processing rote. In general, large grain size of sheet in deep drawing operation cause orange peel effect. Orange peel is a cosmetic defect associated with a rough surface appearance after forming a component from sheet metal. It is called orange peel because the surface has the appearance of the surface of an orange. But smaller grain size results in smoother surface finish.

The effect of “too high” & “too low” blank holder force on deep drawing:

The drawing force required to perform a given operation can be estimated roughly by the formula [1]:

Where P = total punch load, 𝞂_o= average flow stress, D_p = punch diameter, D_o = blank diameter, H = hold-down force, B = force required to bend and restraighten blank, h = wall-thickness, µ = coefficient of friction

If it is too small, wrinkling occurs. If it is too large, it prevents the metal from flowing properly toward the die cavity, resulting in stretching and possible tearing of the sheet metal. Determining the proper holding force involves a delicate balance between these opposing factors [2].

The purpose of the punch and die radius (𝑅_𝑝, 𝑅_𝐷) and the clearance between them:

One of the measures of the severity of a deep drawing operation is the drawing ratio DR. This is most easily defined for a cylindrical shape as the ratio of blank diameter D_b to punch diameter D_p. In equation form,

DR = D_b/

The drawing ratio provides an indication, albeit a crude one, of the severity of a given drawing operation. The greater the ratio, the more severe the operation. An approximate upper limit on the drawing ratio is a value of 2.0. The actual limiting value for a given operation depends on punch and die corner radii (R_p and R_d), friction conditions, depth of draw, and characteristics of the sheet metal (e.g., ductility, degree of directionality of strength properties in the metal)[2].

As discussed earlier the clearance between punch and die ~20 to 40 percent greater than the sheet thickness [1].

Role of Lubrication

Correct lubrication of the sheet metal is essential if friction, wear, and galling are to be held to the lowest possible levels during deep drawing. In fact, deep drawing is impossible if the sheet metal is not lubricated. In actual practice, die materials are selected after trials using one or more candidate production lubricants. If excessive wear or galling occurs, a better lubricant is usually applied. For extremely difficult draws, the best lubricants are usually applied at the outset [5].

References

[1] Dieter, George Ellwood, and D. J. Bacon. 1988. Mechanical metallurgy. London: McGraw-Hill.

[2] Groover, Mikell P., 1939-. Fundamentals of Modern Manufacturing: Materials, Processes, and Systems. Hoboken, NJ:J. Wiley & Sons, 2007.

[3] Askeland, Donald R. 1984. The science and engineering of materials. Monterey, CA: Brooks/Cole Engineering Division.

[4] Callister, William D., and David G. Rethwisch. 2008. Fundamentals of materials science and engineering: an integrated approach. Hoboken, NJ: John Wiley & Sons.

[5] ASM International. 2006. ASM Handbook. Volume 14B, Volume 14B. https://doi.org/10.31399/asm.hb.v14b.9781627081863.

Age hardening of an extremely ductile martensitic steel???

Heat Treatment of Maraging Steels

MARAGING STEELS derive their strength from the formation of a very low-carbon, tough, and ductile iron-nickel martensite, which can be further strengthened by subsequent precipitation of intermetallic compounds during age hardening. The term marage was coined based on the age hardening of the martensitic structure. Maraging steels are highly alloyed low-carbon iron-nickel martensites that possess an excellent combination of strength and toughness superior to that of most carbon-hardened steels (Fig. 1). As such, they constitute an alternative to hardened carbon steels in critical applications where high strength and good toughness and ductility are required. Hardened carbon steels derive their strength from transformation-hardening mechanisms (such as martensite and bainite formation) and the subsequent precipitation of carbides during tempering.

Fig. 1 Strength/toughness combination of 18 Ni maraging steels compared to conventional high-strength carbon steels.

Solution Annealing

The martensitic matrix of maraging steels is prepared for later age hardening through a heat-treating procedure commonly referred to as a solution anneal. Solution annealing entails heating the alloy significantly above the austenite finish (Af) temperature, holding a sufficient time to place the alloying elements in solid solution, and then cooling to room temperature. The most common solution-annealing cycle for the 18 Ni Marage 200, 250, and 300 alloys involves heating to 815 °C (1500 °F) for 1 h followed by air cooling.

Age Hardening

A typical age-hardening treatment after solution annealing usually consists of reheating the alloy into the temperature range of 455 to 510 °C (850 to 950 °F), holding at this temperature for 3 to 12 h, and air cooling to room temperature. In typical treatments at 480 °C (900 °F), the 18 Ni Marage 200, 250, and 300 grades are held 3 to 8 h, whereas the 18 Ni Marage 350 grade is usually held 6 to 12 h at 480 °C (900 °F). The 18 Ni Marage 350 grade can also be aged for 3 to 6 h at 495 to 510 °C (925 to 950 °F). The use of marage steels in applications such as die casting tooling requires the use of an aging temperature of approximately 530 °C (985 °F) to provide an overaged structure that is more thermally stable.

Historical Development

Age hardening of martensite, or maraging, depends on the occurrence of a thermal hysteresis of phase transformations, whereby the reversion of martensite to austenite during reheating occurs at a higher temperature range than the temperature range for martensite formation during cooling. Although knowledge of this thermal hysteresis in Fe-Ni alloys was observed as early as 1927, the first extensive research toward development of commercial iron-nickel maraging steels was conducted in the late 1950s by the International Nickel Company (currently Inco Ltd.). This research culminated in the development of the 20 and 25% Ni maraging steels. In addition to nickel, these two alloys contained 0.3% Al, 1.4% Ti, and 0.4% Nb, which resulted in precipitation hardening of the low-carbon martensitic structure when aged at 425 to 510 °C (800 to 950 °F). Both alloys were reported to exhibit good combinations of strength and ductility at hardness levels of 53 to 56 HRC; however, as reported by Hall, these alloys were abandoned because of their brittleness at extremely high strength levels.

Subsequent work on the iron-nickel system by Decker, Eash, and Goldman revealed that the martensite formed in this binary system could be hardened to appreciable levels through the addition of cobalt and molybdenum. A nickel level of 18% was chosen for this alloy system because nickel levels significantly greater than 18% resulted in the retention of austenite in the as-quenched condition. By the early 1960s, three new maraging steels based on the Fe-18Ni-Co-Mo quaternary alloy system were introduced. These were the 18 Ni Marage 200, 250, and 300 alloys, which are capable of achieving yield strengths of approximately 1380, 1725, and 2000 MPa (200, 250, and 290 ksi), respectively, in combination with excellent ductility and toughness. The nominal compositions of these grades are shown in Table 1. In general, the strength levels attained by these alloys are determined by the cobalt, molybdenum, and titanium contents of each alloy. These three alloys replaced the phased-out 20 and 25% Ni maraging steels.

Table 1 Nominal compositions of standard commercial maraging steels.

Why Tool Steels need a double Tempering Heat Treatment?

 Tempering of Tool Steels

Tempering modifies the properties of quench-hardened tool steels to produce a more desirable combination of strength, hardness, and toughness than obtained in the quenched steel. The as quenched structure of tool steel is a heterogeneous mixture of retained austenite, un tempered martensite, and carbides.

More than one tempering cycle may be necessary to produce an optimum structure. It is normally desirable to transform all retained austenite to ensure complete hardness, improve toughness, and minimize distortion during service. This can be more nearly accomplished by two or more shorter time tempering cycles than by a single and longer cycle.

In the higher-alloy tool steels, a small amount of un tempered martensite is formed from retained austenite during the cool down from the first tempering cycle. It is good practice to double temper to ensure more nearly complete transformation of retained austenite and to temper freshly formed martensite. For some highly alloyed grades of tool steel, triple or quadruple tempering is recommended.

The changes that take place in the microstructure during tempering of hardened tool steels are time-temperature dependent. Time at tempering temperature should not be less than 1 h for any given cycle.

Most manufacturers of high-speed steels recommend multiple tempers of 2 h or more each to attain the desired microstructure and properties. Maintaining recommended tempering times, temperatures, and number of tempers (a minimum of two) ensure attainment of consistent tempered martensitic structures and overcomes uncertainties caused by variations in the amount of retained austenite in the as-quenched condition. These variances are functions of differences in heat chemistry, prior thermal history, hardening temperatures, and quenching conditions. Other factors that influence the tempering requirements of high-speed steels are:

· Increasing the free (matrix) carbon content increases the amount of retained austenite in the as-quenched condition

· The amount of retained austenite significantly affects the rate of transformation, particularly for short tempering cycles. Multiple tempering is more important to attain an acceptable structure if short tempering times are used.

· Cobalt in alloys such as M42 reduces the amount of retained austenite in the as-quenched condition and accelerates the transformation of the retained austenite during tempering.

Enough time should be allowed during tempering for the temperature to be distributed uniformly throughout the tools before time at temperature is counted. This is especially true for low tempering temperatures and for tools that have large sections.

Reference

(ASM hand book of Heat Treating, 1991)

What is Argon oxygen decarburization (AOD) in Secondary Steel Making?

 

Argon Oxygen Decarburization (AOD)

Argon oxygen decarburization offers improved metal cleanliness, which is measured by low unwanted residual element contents and gas contents; this ensures superior mechanical properties. The AOD process is duplexed, with molten metal supplied from a separate melting source to the AOD refining unit (vessel). The source of the molten metal is usually an electric arc furnace or a coreless induction furnace. Foundries and integrated steel mills utilize vessels ranging in nominal capacity from 1 to 160 Mg (1 to 175 tons).

Although the process was initially targeted for stainless steel production, argon oxygen decarburization is used in refining a wide range of alloys, including:

v  Stainless steels

v  Tool steels

v  Silicon (electrical) steels

v  Carbon steels, low-alloy steels, and high-strength low-alloy steels

v  High-temperature alloys and superalloys

Fundamentals

In the AOD process, oxygen, argon, and nitrogen are injected into a molten metal bath through submerged, side-mounted tuyeres. The primary aspect of the AOD process is the shift in the decarburization thermodynamics that is afforded by blowing with mixtures of oxygen and inert gas as opposed to pure oxygen.

To understand the AOD process, it is necessary to examine the thermodynamics governing the reactions that occur in the refining of stainless steel, that is, the relationship among carbon, chromium, chromium oxide (Cr3O4), and carbon monoxide (CO). The overall reaction in the decarburization of chromium-containing steel can be written as:

The equilibrium constant, K, is given by:

Where a and P represent the activity and partial pressure, respectively.

At a given temperature, there is a fixed, limited amount of chromium that can exist in the molten bath that is in equilibrium with carbon. By examining Eq 2, one can see that by reducing the partial pressure of CO, the quantity of chromium that can exist in the molten bath in equilibrium with carbon increases. The partial pressure of CO can be reduced by injecting mixtures of oxygen and inert gas during the decarburization of stainless steel. Figure 1 illustrates the relationship among carbon, chromium, and temperature for a partial pressure of CO equal to 1 and 0.10 atm (1000 and 100 mbar, or 760 and 76 torr). The data shows that diluting the partial pressure of CO allows lower carbon levels to be obtained at higher chromium contents with lower temperatures.

Fig. 1 Carbon-chromium equilibrium curves.

In refining stainless steel, it is generally necessary to decarburize the molten bath to less than 0.05% C. Chromium is quite susceptible to oxidation; therefore, prior to the introduction of the AOD process, decarburization was accomplished by withholding most of the chromium until the bath had been decarburized by oxygen lancing. After the bath was fully decarburized, low-carbon ferrochromium and other low-carbon ferroalloys were added to the melt to meet chemical specifications.

Dilution of the partial pressure of CO allows the removal of carbon to low levels without excessive chromium oxidation. This practice enables the use of high-carbon ferroalloys in the charge mix, avoiding the substantially more expensive low carbon ferroalloys.

Fig. 2 Composition changes in refining type 304-L stainless steel using electric arc furnace practice and argon oxygen decarburization

Equipment

The processing vessel consists of a refractory-lined steel shell mounted on a tiltable trunnion ring (Fig. 2). With a removable, conical cover in place, the vessel outline is sometimes described as pear shaped. Several basic refractory types and various quality levels of the refractories have gained widespread acceptance (Ref 4, 5). Dolomite refractories are used in most AOD installations; magnesite chromium refractories are predominant in small (<9 Mg, or 10 ton) installations.

Fig. 2 Schematic of argon oxygen decarburization vessel.

Processing

1.      Stainless Steels

Charge materials (scrap and ferroalloys) are melted in the melting furnace. The charge is usually melted with the chromium, nickel, and manganese concentrations at midrange specifications. The carbon content at meltdown can vary from 0.50 to 3.0%, depending on the scrap content of the charge. Once the charge is melted down, the heat is tapped, and the slag is removed and weighed prior to charging the AOD vessel.

In the refining of stainless steel grades, oxygen and inert gas are injected into the bath in a stepwise manner. The ratio of oxygen to inert gas injected decreases (3:1, 1:1, 1:3) as the carbon level decreases. Once the aim carbon level is obtained, a reduction mix (silicon, aluminum, and lime) is added. If extra-low sulfur levels are desired, a second desulfurization can be added. Both of these steps are followed by an argon stir. After reduction, a complete chemistry sample is usually taken and trim additions made following analysis.

Carbon and Low-Alloy Steels

The refining of carbon and low-alloy steels involves a two-step practice: a carbon removal step, followed by a reduction/heating step. The lower alloy content of these steels eliminates the need for injecting less than a 3:1 ratio of oxygen to inert gas. Once the aim carbon level is obtained, carbon steels are processed similarly to stainless steels. Figure 5 illustrates carbon content and temperature relationship for the AOD refining of carbon and low-alloy steels. Because the alloy content of these grades of steel is substantially lower than that of stainless steel and because the final carbon levels are generally higher, there is no thermodynamic or practical reason for using an oxygen, inert-gas ratio of less than 3:1.

Oxidation measurements indicate that all of the oxygen reacts with the bath and that none leaves the vessel unreacted. By monitoring and recording the oxygen consumption during refining, very close control of end point carbon is achieved. Because the oxygen and inert gases are introduced below the bath and at sonic velocities, there is excellent bath mixing and intimate slag/metal contact. As a result, the reaction kinetics of all chemical processes that take place within the vessel are greatly improved.

Decarburization

In both stainless and low-alloy steels, the dilution of oxygen with inert gas results in increased carbon removal efficiencies without excessive metallic oxidation. In stainless grades, carbon levels of 0.01% are readily obtained.

Chemistry Control

The excellent compositional control of AOD-refined steel is indicated in Table 2 for a ten-heat series of high-strength low-alloy steel. The injection of a known quantity of oxygen with a predetermined bath weight enables the steelmaker to obtain very tight chemical specifications.

Nitrogen Control

    Degassing in argon oxygen decarburization is achieved by inert gas sparging. Each argon and CO bubble leaving the bath removes a small amount of dissolved nitrogen and hydrogen. Final nitrogen content can be accurately controlled by substituting nitrogen for argon during refining. Nitrogen levels as low as 25 to 30 ppm can be obtained in carbon and low-alloy steels, and 100 to 150 ppm N can be obtained in stainless steels. The ability to obtain aim nitrogen levels substantially reduces the need to use nitrided ferroalloys for alloy specification, and this also minimizes the use of argon. Hydrogen levels as low as 1.5 ppm can be obtained.