Fatigue Properties of Modified Steels Carburized at High Temperatures

The objective of phase XI is to compare the performance of carburized steel in bending and axial fatigue loading conditions. Carburization is a surface heat treatment process which adds carbon to the surface of a material to increase its hardness and other mechanical properties. Such heat treatments are typically performed on materials used in shafts and gear applications. SAE 8615 was chosen for this study as it is a common carburizing grade with good toughness.

A total of ten iterations of constant amplitude and overload axial and 4-point bending fatigue were performed for three steel conditions as shown in Table 1.

Table 1: Phase XI – Testing Methods

table1.png

* Simulated carburized core is defined as through carburized.

** Shallow case depth is defined as 10% of the gage radius is carburized.

*** Deep case depth is defined as 20% of gage radius is carburized.

The constant-amplitude axial fatigue tests provide information about the cyclic and fatigue behavior of a material.

Background on SAE 8615 Steel:

The SAE 8615 is principally a carburizing grade of steel. It finds medium-strength applications where high surface hardness and high core toughness are required. This alloy is used as forged shafts and bearings in automotive and aerospace industries.

  • Machinability: SAE 8615 is readily machinable following suitable cooling from forging or subsequent heat treatment.
  • Weldability: This alloy may be welded by normal fusion methods prior to carburizing, hardening or tempering.

 

Monotonic Tension Tests:

Two monotonic tension tests are conducted to ensure accuracy of the results. The engineering stress-strain and true stress-strain curves are obtained from tension tests (Fig. 1 and Fig. 2) from which various properties, as given in Table 2, are calculated.

Fig1

Figure 1: Engineering Stress – Strain Curves

fig2.png

Figure 2: True Stress – Strain Curves

 

Table 2: Monotonic Tensile Properties

table2.png

Cyclic Tests:

When a material is loaded in a cyclic mannerit first shows a transient response for a number of loading cycles until stabilizing for a time-constant response. This phenomenon is called cyclic stabilization.

In this first of four blogs, we will evaluate the response of SAE 8615 steel in first of three conditions – the simulated carburized core condition. Fourteen constant-amplitude strain-controlled axial fatigue tests were conducted on the heat treated coupons, which were loaded at 7 different strain amplitudes ranging from 0.2% to 2.0%. The displacements were measured using an extensometer and then converted to strain values. The plot of stress amplitude vs reversals to failure obtained from these tests tells us if a material cyclically hardens or cyclically softens. Figure 3 shows a comparison between the monotonic tensile curve and the cyclic stabilized fatigue curve, and indicates SAE 8615 cyclically softens.

fig3.png

Figure 3: Composite Plot of Monotonic and Cyclic Stress-Strain Curves

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Figure 4: True Stress Amplitude vs Number of Cycles

For any given strain amplitude the true stress amplitude drops with number of cycles until the material finally fails (denoted by straight vertical lines), as shown in Figure 4. Two exceptions are for strain amplitudes of 0.25% and 0.3%, where the stress amplitude is 500 MPa and 560 MPa respectively. At these strain amplitudes, constant stress amplitude is developed and the samples do not fail when cycled up to 1×106 cycles. This is the fatigue strength or endurance limit of the material.

Once the cyclic stabilization occurs, the cyclic properties of a material are determined. The cyclical properties for SAE 8615 steel in simulated carburized core condition are shown in Table 3. These properties will be useful to engineers when they are designing a component for fatigue.

Table 3: Cyclical Properties of SAE 8615 in Simulated Carburized Core Condition

table3.png

It is also interesting to analyze is the strain-life curve of the material in constant-amplitude axial fatigue tests. As shown in Figure 5, the green dots and the corresponding curve show the true strain amplitude vs reversals to failure of SAE 8615 steel in simulated carburized core condition. ‘Reversals to failure’ is a term used in fatigue instead of cycles; 1 cycle = 2 reversals. From Figure 5, we can determine the fatigue life of a material.

fig5.png

Figure 5: True Strain amplitude vs Reversals to Failure

Bending Fatigue Tests:

4-point bending fatigue tests were conducted after axial fatigue testing. The true strain amplitude vs reversals to failure results are shown in Figure 6.

fig6.png

Figure 6: True Strain Amplitude vs Reversals to Failure in a 4-point Bending Test

Though a curve trend is not present in the above graph, a select point comparison between the curves from axial and 4-point bending tests can be extrapolated, as shown in Table 4.

Table 4: Fatigue Life Comparison of Two Cases

table4.png

As seen from the results above, the SAE 8615 shows similar behavior in both constant-amplitude axial and 4-point bending fatigue testing. The axial fatigue curve is more linear than the four-point bending fatigue curve where the strain amplitude at both 1×104 and 1×105 cycles is lower in four-point bending fatigue than in axial fatigue.

This blog reviewed the axial and four-point bending fatigue of SAE 8615 steel in simulated carburized core condition. The next blogs in this series will review the SAE 8615 steel performance in shallow- and deep-case depth conditions in axial, overload axial, 4-point bending and overload 4-point bending conditions.

Stay tuned for some really interesting fatigue life analyses!

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Neuber Stress Plots

This article explains how a “Neuber Stress” is calculated from fatigue test data obtained from un-notched axially-loaded specimen.

In the presence of a geometric stress concentration feature on a part, the fatigue life of that part is often under-predicted as a result of over-prediction of the effect of stress concentration. Actually, the stress around these zones will be reduced by plastic redistribution of stresses by the material. The Neuber Stress correction factor is used to reduce the apparent high stresses caused near geometric stress concentration zones. Traditionally, when a part with a geometric stress concentration is modelled with a linear elastic material model, the geometric stress concentration factor Kt, is used to evaluate stresses near the stress concentrations, resulting in, high stress values.  Being a conservative design strategy, this leads to a potentially heavier component. The linear elastic material model does not account for material plasticity near the stress concentration zones. A Kt value of 3 or higher would be associated with traditionally found stress concentration zones, resulting in a three-fold higher stress value. Conventionally, the stresses around stress concentration zones are calculated by multiplying the nominal stress Sn by ‘kt’, as shown in Figure 1. In FEA packages, if the same linear elastic material models are used for stress analyses, a maximum stress result of a higher magnitude will be observed, as shown in Figure 2.

Fig.1

Figure 1[1]: Conventional method of calculating stresses around stress concentration zones.

Fig.2.PNG

Figure 2[1]: FEA analysis calculating stress at a change in geometry using an elastic model (transition radius)

So, how is the actual stress around the stress raisers attained?

In Figure 3, the linear elastic model predicts a stress SLE and a corresponding strain value of eLE. To find the corrected stress and strain values, a curve of constant total energy (usually a hyperbola) is used, and is intersected with the elasto-plastic stress strain curve of the material. The point where this intersection occurs, gives the corrected stress and strain values, Sc and ec respectively.

Fig.3

Figure 3[2]: Stress correction considering the material plasticity.

The corrected strain value, ec, is higher than the modelled strain, eLE, which occurs as a result of the plastic redistribution of material around the stress concentration zones.

For factoring the correction in, the total energy in both cases is equated and the corrected stress and strain values are determined i.e., SLE * eLE =  Sc * ec .

For static analyses, a spreadsheet of Neuber (corrected) stress values can be found via this link: Static Analysis Neuber Corrected Stress Values.

In case of a fatigue analysis, the procedure remains fairly similar. Here, the material behavior is divided into two parts: the material follows a cyclic stress-strain curve for initial loading, and it follows a hysteresis stress-strain behavior for subsequent loading reversals. This is illustrated in Figure 4.

Fig.4.PNG

Figure 4[1]: Stress correction in material fatigue cases

Note: The graph in Figure 4 uses ‘σ’ and ‘ϵ’ for SLE  and eLE , respectively, as discussed in this blog.

A Neuber stress plot, is essentially a plot of energy vs fatigue life. The “energy” here, is calculated as √(E ∗ Δe ∗ ΔS), E being the Modulus of Elasticity of the material. This quantity, having the units of stress, represents the energy at a fatigue hot-spot, such as the region at the root of a stress concentration, a notch for example, from where a fatigue crack can initiate.

The Sc * e= SLE * eLE equation can be written as:

Sc * e* E= SLE * eLE  * E

Sc * ec   * E = SLE * SLE

√(Sc * ec * E) = SLE  = Neuber stress = S.

After running fully axial (reversed tension/compression) fatigue tests, the constant amplitude stress/strain data will be generated.  The elastic modulus, and the stress and strain ranges are taken from the stabilized stress-strain plots, and the Neuber Stress is computed to create a fatigue life plot, as shown in Figure 5.

Fig.5.PNG

Figure 5[1]: Axial test specimen, stabilized stress-strain plots.

The Neuber stress, √(E ∗ Δe ∗ ΔS) is then plotted along Y-axis to generate a Neuber Stress Plot, and ultimately the fatigue life of a specimen under a given Neuber Stress value is estimated. The material plasticity correction automatically gets built into the graph. An example of such a graph is shown in Figure 6.

Fig.6.PNG

Figure 6[1]: An example of a  Neuber Stress Plot

The advantage of a Neuber stress plot over a stress-life or strain-life plot is that by integrating three variables – stress, strain and the elastic modulus of a material, it accounts for the plastic redistribution of stresses around the stress concentration zones and models it more accurately.

Neuber stress plots can be useful to fatigue design engineers, and part 2 of this blog post will delve more into this.

REFERENCES:

  1. Al Conle’s study “Plasticity Corrections for Elastic Analysis Results: Neuber Method –http://fde.uwaterloo.ca/Fde/Notches.new/neuberStress4AISIpt1.pdf
  2. Neuber Stress explanation from Abbott Aerospace – https://www.abbottaerospace.com/neuber-method-for-reducing-elastic-stress-values
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Mechanical Properties versus Carburization Depth

In this article we will examine the relationship between ultimate tensile strength, fatigue strength, and elongation versus the depth of carburization. The data for this comparison was taken from several iterations discussed in previous articles and shown in Table 1. Depending on the data set, the depth of carburization (case depth) is defined as the distance from the surface to 50 or 58 HRC, expressed as a percent of the 0.200 inch gage diameter.

Table 1

Table 1

The first set of data is for the simulated carburized steel core samples (Table 1: Carb Core) with a hardness of approximately 45 HRC. This data set is made up of iteration #39 (SAE 8620 steel), iteration #82 (SAE 86B20 steel), and iteration #130 (DIN 20MnCr5 steel). These were the only three sample iterations with this core hardness, which was chosen to be consistent with the next data set.

The second data set is the simulated case-core composite samples (Table 1: Light Case) from a previous article (Carburized Case-Core Samples). The iteration numbers were #62 (SAE 8620 steel), #63 (4320 steel), and #70 (8620 steel). These samples were carburized with an effective case depth to 50 HRC that was 5% (#63) or 10% (#62, #70) of the diameter. The core was hardened to approximately 45 HRC.

The third data set was the early through-carburized samples (Table 1: Thru-Carb Early) that were typically ground and polished after heat treatment. These were iteration numbers #38, #48, #50, #54, #56, #58, #60, and #71. The steels were made up of SAE grades 8620, 4620, 4320, 5120, 9310, and 8822. These samples were carburized at 927 °C for 24 to 26 hours. The surface hardness was typically above 60 HRC and the core hardness was above 55 HRC. The depth of carburization to 58 HRC was approximately 37.5% of the diameter.

The fourth data set was the through-carburized samples (Table 1: Thru-Carb Late) that were not ground after heat treatment. These were iteration numbers #73, #75, #79, #83, #87, #88, #89, #90, #95, and #137. The SAE grades used were 41B17, 86B20, 8620, and DIN 20MnCr5. The heat treatment cycle, hardness and carburization depth were similar to the third data set.

The last data set was the through-carburized samples (Table 1: Actual Thru-Carb) that were carburized completely to the center of the bar. These were iterations #40, #41, and #141. The first two iterations were SAE 8695 steel and the third was DIN 20MnCr5 steel carburized at 927 °C for 36 hours.

A graph of the ultimate tensile strength and fatigue strength versus the carburized depth is shown in Figure 1. By carburizing to 5 to 10% of the gauge diameter, the average ultimate strength increases from 1507 MPa with no case to 1715 MPa. By increasing the case to approximately 37.5% of the diameter, the average ultimate strength decreases to 1667 MPa for the samples ground and polished after heat treatment, and 1525 MPa for those not ground and polished after heat treatment. This indicates there is a benefit to grinding and polishing after heat treatment by removing intergranular oxidation. By completely through carburizing the gage diameter, the average ultimate strength further decreases to 1147 MPa. For these axial test bars, a small amount of high carbon case depth is beneficial for ultimate strength. However, increasing the carburization depth too much can significantly decrease the ultimate strength, which supports the practice in industry to develop components with a shallow case depth to increase properties.

Figure 1 also shows as the depth of carburization increases so does the spread between the minimum and maximum tensile and fatigue strength. However, data sets three and four, with the largest spread, had eight and ten samples respectively, while all other data sets had only three samples. More data would be necessary to confirm this observation. The fatigue strength follows the same trend as the ultimate strength, and the amount of spread appears proportional.

SMDI Blog 40 Figure 1
Figure 1

A graph of the elongation versus the carburized depth is shown in Figure 2. The core samples with no carburized case depth (Carb Core) have the highest average elongation at 29.6%. Carburizing to 5 to 10% of the diameter decreased the average elongation to 16.2% (Light Case). Carburizing to 37.5% of the diameter further decreases the average elongation to 3.7% (Thru-Carb Early) and 2.4% (Thru-Carb Late). Through carburizing the samples to 50% of the diameter resulted in an average elongation of 1.7%. The maximum amount of spread was in the core samples at the far left side of the graph.

SMDI Blog 40 Figure 2
Figure 2

In conclusion, carburizing these axial test bars through the diameter provides the lowest ultimate and fatigue strength, and the lowest ductility. The maximum combination of ultimate strength, fatigue strength, and ductility is obtained by carburizing to 5 to 10% of the gage length diameter; further validating industry practice.

 

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Through Carburized 20MoCr4 Steel, 26-Hour Atmosphere Cycle versus 36-Hour Vacuum Cycle

In this article we will compare two through-carburized 20MoCr4 steel iterations. The composition of the 20MoCr4 steel is shown below in Table 1. Iteration #137 was atmosphere through carburized using the traditional 927 °C, twenty-six hour cycle which provides a surface hardness of 61 to 64 HRC and a core hardness of 55 to 60 HRC. This cycle does not provide a uniform carbon content of 0.80 to 0.90 weight percent through the 0.200 inch (5.08 mm) gage length diameter. Judging by the hardness, the core is approximately 0.42 weight percent carbon. Iteration #141 was vacuum through carburized using a thirty-six hour cycle at 927 °C, which does provide uniform carbon content and a uniform hardness through the entire cross section.

Blog 38c - Table 1

The mechanical properties, fatigue strength, and hardness for iterations #137 and #141 are shown in Table 2. The most significant difference is the strength. The yield and ultimate strengths of iteration #137 (1,265 and 1,394 MPa respectively) are typical of what we have seen in the past few articles with through-carburized samples that have not been ground after heat treatment. However, the yield and ultimate strengths of iteration #141 (965 and 965 MPa respectively) are significantly lower than iteration #137 and are the same indicating there is no plastic deformation prior to fracture. Essentially the sample is weaker and brittle. The fatigue strength (461 MPa for iteration #137, and 465 MPa for iteration #141) was not affected by the yield/ultimate strength differences as both iterations are nearly identical.

Blog 38c - Table 2

The potential causes for the decrease in strength would be; atmosphere versus vacuum carburizing, an increase in grain size as a result of the longer cycle, or the deeper carbon penetration.  In the prior article, “Atmosphere versus Vacuum Through-Carburized Boron Steels” we concluded there was no significant difference in performance between atmosphere and vacuum carburizing. The data is shown in Table 3.

Blog 38c - Table 3

The microstructure photos in the AISI database for iterations #137 and #141 were examined. The grain size appears to be the same for both and no abnormal growth was noted. The typical grain diameter is less than 20 microns. From this it does not appear grain size is responsible for the difference in strength. This leaves depth of carburization as the cause.

Iterations #40 and #41 were examined in several prior blog articles. These two iterations were made from vacuum melted 100 pound heats of 8695 steel. The steel was cast and forged into bar, from which test bars were machined. The carbon content was relatively constant through the cross section, much the same as iteration #141 in this article. Iteration #40 had no intergranular oxidation (IGO) at the surface, while iteration #41 had IGO at the surface. Both iterations were carburized using a normal industry cycle, however iteration #40 was copper plated prior to carburizing to prevent IGO. The data is shown in Table 4. Iteration #41 has similar strength as iteration #141 in this article (979 MPa and 965 MPa respectively). From this it appears the depth of carburization is the likely cause of the lower strength of iteration #141 compared to #137. However, there is some room for doubt as the ultimate strength of iteration #40 is higher than #41 (1,496 MPa and 979 MPa respectively) and no cause has been determined. More work would be necessary to confirm this conclusion.

Blog 38c - Table 4.png

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The Effect of Bainite on Strength and Fatigue in Through-Carburized SAE 8620 Test Bars

A case microstructure may consist of martensite, retained austenite and bainite, and depends upon alloy and cooling rate. The impact of bainite on fatigue is not well known.  Sometimes carburized parts are intentionally austempered to increase the amount of bainite and achieve higher toughness. How does the amount of bainite affect strength and fatigue? This article will attempt to answer this question.

The SAE 8620 test bars in this experiment were through-carburized at 927 C for 26 hours. This heat-treat cycle does not completely through-carburize the 0.200 inch gage diameter, but typically leaves the core hardness at or above 55 HRC. All test bars were finished machined and polished prior to heat treatment. After heat treatment and prior to testing, a final polish using 600 grit emery paper was done. This would remove some or all of the intergranular oxidation (IGO).

The percent of bainite was controlled by isothermally quenching the test pieces at an elevated temperature.

  • Iteration #87 was direct quenched in cold oil after carburizing to achieve 100 percent martensite.
  • Iteration #88, with 25 percent bainite and 75 percent martensite, was reheated to 857 °C for 30 minutes and quenched to 232 °C and held for 30 minutes.
  • Iteration #89, with 50 percent bainite and 50 percent martensite, was reheated to 857 °C for 30 minutes and quenched to 232 °C and held for 50 minutes.
  • Iteration #90, with 75 percent bainite and 25 percent martensite, was reheated to 857 °C for 30 minutes and quenched to 232 °C and held for 150 minutes.
  • Iteration #95, with 100 percent bainite, was reheated to 857 °C and quenched to 316 °C and held for 240 minutes.

After heat treatment all iterations were metallurgically evaluated to confirm the amount of bainite predicted in the TTT diagram.

The mechanical properties, fatigue strength, and hardness are shown in Table 1. A graphical representation of strength versus the percentage of bainite is shown in Figure 1. From 0 to 50 percent bainite the ultimate strength is constant at around 1400 MPa. At 75 to 100 percent bainite the ultimate strength increases to around 1800 MPa. From 0 to 50 percent bainite the yield strength decreases slightly from 1,252 MPa to 1,114 MPa. Above 50 percent bainite the yield strength increases to a maximum 1,556 MPa at 100 percent bainite. The highest combination of yield strength and ultimate strength is at 100 percent bainite.

Table 1Table 1

Figure 1Figure 1

A graph of percent elongation versus the percent of bainite is shown in Figure 2. From 0 to 75 percent bainite the elongation gradually increases from 1 to 1.3 percent. From 75 to 100 percent bainite the elongation dramatically increases to 15 percent. The highest combination of yield, ultimate strength and elongation is at 100 percent bainite.

Figure 2Figure 2

A graph of fatigue strength versus percentage bainite is shown in Figure 3. There is no clear correlation as the maximum fatigue strength of 758 MPa is at 25 percent bainite. The fatigue strength of all other iterations is relatively constant at just below 600 MPa. The difference in fatigue strength is likely due to normal variation.

Figure 3Figure 3

A graph of surface hardness versus percentage bainite is shown in Figure 4. From 0 to 50 percent bainite the surface hardness is relatively constant at around 700 Brinell or 63 HRC. At 75 percent bainite the surface hardness decreases to 653 Brinell or 60 HRC, and at 100 percent bainite the surface hardness further decreases to 514 Brinell or 52 HRC.

Figure 4

Figure 4

A graph of ultimate strength and fatigue strength versus hardness is shown in Figure 5. From 514 to 653 Brinell the ultimate strength increases slightly from 1,785 MPa to 1,883 MPa. Above 653 Brinell the ultimate strength decreases rapidly to as low as 1,390 MPa. This is a documented relationship that was presented in the prior blog article, “Hardness versus Strength.” Ultimate strength increases linearly with hardness to about 550 Brinell and then decreases rapidly above that. From 514 to 682 Brinell the fatigue strength is relatively constant at around 600 MPa. At 710 Brinell fatigue strength increases to 758 MPa. It is tempting to conclude the increase in fatigue strength is related to the increase in hardness. However, in the previous blog article, “Hardness versus Fatigue Strength” it was shown that fatigue strength is largely a percent of the ultimate strength. It must be pointed out that there is a considerable amount of scatter or variation to the data in this relationship. It is believed the difference, and lack of difference, in fatigue strength observed in the current study is normal variation and not related to any change in hardness or microstructure.

The optimum condition for maximizing yield and ultimate strength, as well as ductility and fatigue strength is the 100 percent bainite condition. However, this condition also comes with a lower than normal hardness for most carburized components. This would likely reduce contact strength and fatigue, which must be considered in a component such as a gear.

Figure 5Figure 5

 

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Atmosphere versus Vacuum Through-Carburized Boron Steels

In this article we will compare two through-carburized boron steels which were both atmosphere and vacuum heat treated. As in previous axial fatigue tests; approximately 30 test samples were prepared to ensure there were three replicate tests at a minimum of 6 strain levels to develop a complete fatigue curve. These test bars were finished prior to the carburizing heat treatment with no subsequent grinding. The through-carburizing cycle of 26 hours at 927 °C does not completely through carburize the 0.2 inch gage length diameter, but typically leaves the core hardness at or above 55 HRC. The two steel grades are SAE 41B17 and SAE 86B20. The chemical composition of each is shown in Table 1. The primary difference between the two steels is the nickel content, which is significantly lower in the SAE 41B17 grade as it is a residual element.

Table 1
Table 1

The mechanical properties, fatigue strength and hardness are shown in Table 2. Iterations #73 and #75 were atmosphere carburized while iterations #79 and #83 were vacuum carburized. The elongation and reduction of area is 1% or less for all iterations with the exception of iteration #79. This iteration has the lowest surface (case) hardness at 653 Brinell or 60 HRC. The remaining iterations are 665-710 Brinell or 61-63 HRC. All of these hardness values, including iteration #79, would be acceptable for a typical carburized component.

Table 2Table 2

A graphical representation of the yield and ultimate strength, as well as the fatigue strength, is shown in Figure 1. The fatigue strength is the stress level which provides one million cycles during strain controlled axial fatigue testing. Iteration #79 has the lowest hardness, however, the yield and ultimate strength is approximately 40% greater than the others. This indicates a lower hardness may be beneficial in improving fatigue properties. While the fatigue strength does vary from 235-536 MPa for these tests there is no clear cause (chemistry or carburizing method) for the variation and is considered to be within normal fatigue result variation.

Figure 1 - New
Figure 1

Plots of surface hardness versus ultimate strength and hardness versus fatigue strength are shown in Figure 2. The plot shows no correlation between hardness and fatigue strength. This was also observed in the prior blog article, “Strength and Fatigue Life versus Carbon Content at High Hardness (60 HRC).” Obviously, there is a relationship between hardness and fatigue strength (see previous blog article, “Hardness versus Fatigue Strength”) but it has considerable variability and is not evident here. Figure 2 shows a significant increase in ultimate strength when the hardness decreases to 653 Brinell. This is in agreement with the previous blog article, “Hardness versus Strength,” where strength increased linearly with hardness to just below 600 Brinell and then decreased rapidly with any further increase in hardness. SAE J413 teaches us strength should increase with hardness beyond 600 Brinell. However, this does not appear to be true with high carbon carburized steel.

Figure 2 - NewFigure 2

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A Look Ahead at Bar Fatigue in 2018

The SMDI Bar Steel Fatigue subcommittee is celebrating 20 years of working together to provide the industry with much needed fatigue data for critical applications. Past work includes the completion of over 200 iterations; 164 released to the public. This work continues in 2018 by working on four studies for key ground vehicle OEMs and suppliers that comprise the subcommittee. The subcommittee concentrates test efforts on those steel grades and process combinations used in safety critical components. Test plans for 2018 include:

  1. Completion of Phase XIII; rotating bending fatigue tests on both SAE 8620 and SAE 9310 steel. The tests results will be compared to previous axial fatigue results for the same steels.
  2. Completion of Phase XIV; SAE 9310 gas carburized to both shallow (5% of bar diameter) and deep (10% of bar diameter) case depths.
  3. Initiation of Phase XVI by testing a SAE 4120 steel modified to be vacuum carburized at high temperatures up to 1900 °F. The test results will be compared to previous iterations of the same grade (not modified) to determine if the high vacuum carburizing temperature and accelerated cycle time is an acceptable process alternative.
  4. Initiation of Phase XVII by testing SAE 8620 steel at elevated temperatures to determine long term durability within harsh environments such as engines and transmissions. The material will be tested at room temperature, 350 °F and 675 °F. The test results will be used to validate formulas currently used in the industry to estimate elevated temperature performance.

Further blog discussions on all of these test phases will be provided when the data becomes available to the public. If you are interested in these results earlier, committee membership is available (contact danderson@steel.org).

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2017 Bar Fatigue Blog – Year in Review

During 2017 the SMDI Bar Steel Fatigue sub-committee continued work on numerous studies, test protocols, as requested by the committee members who represent key ground vehicle OEMs and suppliers. The committee continues to concentrate test efforts on those steel grades and process combinations used in safety critical components. Accomplishments in 2017 include:

  1. Phase XI of the test protocol was completed. The material tested in this phase was SAE 8620 with carburized case depths of 5% (shallow) or 10% (deep). The material was subjected to the standard axial fatigue tests and additionally to four-point bending fatigue tests to compare and correlate results. As in previous phases both the axial and bending tests included overload stress conditions. As part of the study, members of the committee also provided residual stress measurements of selected samples.
  2. The team also completed Test Phase XV. DIN 16MnCr5 steel at carburized case depths of 5% and 10% was evaluated. As in Phase XI the material was tested in both axial and bending fatigue to compare and correlate results. As part of the study a member of the committee completed failed sample fracture analysis to track the location and cause of fatigue failures.
  3. Rotating bending fatigue tests were completed on SAE 8620 and DIN 20MnCr5 steel as part of Test Phase XIII. These tests were run to compare and correlate results to previous axial fatigue results.
  4. The sixth update of the on-line publically available database to 164 steel/process iterations. An additional 30 iterations are being held confidentially to the committee members who fund the research. These iterations represent 13 completed test protocols.

Further blog discussions on all of these test phases will be completed at a later date when the data becomes available to the public. If you are interested in these results earlier, consider becoming a committee member. For more information visit www.autosteel.org or contact David Anderson.

image

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Carburized Case-Core Samples

In this article we will examine Iterations #62, #63 and #70. These are carburized samples with a case/core composite to simulate an actual carburized part. Until now, all of the carburized samples we have looked at have been through carburized or nearly through carburized. The intent is to provide fatigue properties of the carburized case (through carburized) and for the noncarburized core region (via quench and temper heat treatment) to predict fatigue life of a case/core composite. This assumes the microstructure of the case is homogeneous as is the microstructure of the core.

The results for the case-core samples are shown in the table below, along with a summary of the previous through carburized data. Two of the case-core iterations were made from different heats of SAE 8620 steel (iterations #62 and #70), and one was made from SAE 4320 steel (iteration #63). The effective case depths were 0.024 inches (20%) for iteration #62, 0.011 inches (10%) for iteration #63 and 0.022 inches (20%) for iteration #70. The average ultimate strength is 1715 MPa, and the average yield strength is 1368 MPa. These results are high and they are consistent, compared to the through-carburized data we have previously viewed. In addition, the elongation, and reduction of area are also high, compared to the through-carburized samples. The fatigue strength (strength at 10^6 cycles) is also higher and more consistent than we have seen before. The fatigue ratio is about the same, however the consistency is better. The average surface hardness for the case-core samples is lower than either of the through-carburized groups. The case-core samples were ground after carburizing, which is the reason for the lower hardness. The carburized case was relatively thin, and the grinding process removed the hard outer layer. These samples are below the minimum surface hardness specification of 57 or 58 HRC, normally used with carburizing. The core hardness of the case-core samples is significantly lower than the through-carburized samples, which is expected due to the nature of the heat treatment. However, the core hardness is at the upper end of what is typical of a carburized component. It is possible the core hardness may have an effect on fatigue strength, as the sample with the highest core hardness has the greatest ultimate tensile strength. However, with only three samples this cannot be confirmed. The case depth, within the range used, does not appear to have any significant effect on strength. Again, more test results are necessary to make a firm conclusion.

SMDI Blog 39 Table

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Mechanical Properties and Fatigue Strength of the Early Through-Carburized Samples

In this article we will examine the mechanical properties and fatigue strength of the early through-carburized iterations in the SMDI fatigue program. During this time period the test bars were typically ground and polished after carburizing in order to eliminate any distortion as a result of the heat treatment. The through-carburizing heat treatment was a 24-26 hour cycle at 1700 oF (927 oC). This cycle did not completely through carburize the 0.200 inch (5.08 mm) gage diameter to a consistent high carbon level but typically left the core hardness at or above 55 HRC.

The mechanical properties and the fatigue strength, defined as the stress at one million cycles, are shown in Table 1. Additional information in the table includes: the ratio of the fatigue strength to the ultimate tensile strength; surface and core hardness; and the laboratory where samples were machined and tested. Most of the elongation values are less than 2% which is reasonable for these high hardness, high carbon iterations. However, the elongation of iterations 48 and 54 was significantly higher at 12.5% and 7.7% respectively. In addition, the ultimate tensile strength of iteration 48 was considerably greater than any of the others at 2,227 MPa. Iterations 48 and 54 were both tested at the University of Toledo while most of the others were tested at the University of Waterloo. Both of these iterations were made from 4620 steel, which as a medium nickel steel (1.8% nickel) is expected to have increased ductility. However, the 4320 steel is also a medium nickel steel and 9310 is a high nickel steel (3.3% nickel) and neither of these had ductility comparable to the 4620.

Metallurgically speaking the surface hardness of iteration 48 is lower than most other samples except for iteration 50 which is the 4320 sample. This could indicate an elevated tempering temperature was used on iteration 48 after carburization which would tend to increase the ductility. However, this would not explain the high elongation on iteration 54 and the low elongation on iteration 50. The cause of the high strength and high ductility of the two 4620 samples could be related to the amount of stock that was removed from the surface during the grinding process. If more stock was removed from the surface of these two samples the carbon content at the final test specimen surface would be reduced which would increase the strength and ductility. It is also possible these two iterations were simply carburized with a lower carbon potential.

SMDI Blog 37 Table 1
Table 1

Figure 1 provides a graphical representation of the ultimate tensile strength data. It is obvious the first 4620 iteration (48) has an ultimate tensile strength which is greater than any of the other samples. The first three iterations on the left of the chart were not ground, but only polished after carburizing.  These iterations averaged about 1200 MPa in ultimate tensile strength. The tensile strength of the remaining iterations, without the 1117 steel, averaged about 1700 MPa. If we ignore the two 4620 iterations with high elongation the average ultimate strength is still above 1600 MPa. This indicates grinding after carburizing has a beneficial effect on the test specimen strength. The 1117 iteration (52) had a lower strength level which is likely a result of the high level of manganese sulfide inclusions as well as the fact this was a plain carbon steel with no alloy content.

The two 8695 samples are the best representation of a through-carburized material since they were produced with 0.95% carbon through the cross section. They were produced as 100 pound vacuum melted heats which were forged into test samples then machined and carburized. The carburization was not used to introduce any more carbon but rather to introduce intergranular oxidation (IGO) at the surface. The first 8695 iteration (40) was copper plated prior to carburizing to prevent IGO at the surface. Without IGO the ultimate tensile strength is around 1500 MPa, and with IGO it is around 1000 MPa. This difference may be related to the IGO or possibly the carburization heat treatment. The alloy content of the different steels does not appear to make much difference in strength, except for the 1117 steel. The 9310 steel has the highest alloy content followed by the 4320 and then the 4620. These grades do not have a higher ultimate tensile strength than the more common grades such as 8620, 5120 and 8822.

SMDI Blog 37 Figure 1Figure 1

Figure 2 is a graphical representation of the fatigue strength data. The 4320 iteration (50) had the highest fatigue strength, while the lowest was the 1117 iteration (52). The first 8695 iteration (40), without the IGO, had a lower fatigue strength than the second one (41), with the IGO. This indicates fatigue strength does not necessarily correlate with the ultimate strength and in this case the IGO does not appear to be detrimental toward fatigue life. However, the first three iterations, along with the 1117 iteration, have lower fatigue strengths than the other iterations and do correlate to the ultimate strength.

 

SMDI Blog 37 Figure 2.jpgFigure 2

 

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