Fabrication of Cu-Infiltrated Journal Bearing by Binder Jetting Additive Manufacturing

Abstract

In this study, considering the economic feasibility of products that can be produced through the binder jetting additive manufacturing process, 316L stainless steel, a widely used material with a wide particle size ranging from 15 to 106 μm, was used. The lubrication effect was increased by internal patterning through design for additive manufacturing, and journal bearing parts with excellent load resistance and wear resistance were implemented by using wear-resistant Cu as an infiltration material. In addition, to investigate the amount of Cu infiltrated as a function of porosity, the parts were pre-sintered from 1423 K to 1573 K, and the best performance was obtained when Cu was infiltrated after pre-sintering at 1473 K. As a result of rig testing of Cu-infiltrated journal bearings, mechanical properties were obtained that were more than 50% improved compared to those of mass products.

1. Introduction

The binder jetting (BJT) additive manufacturing (AM) process has the advantages of material selectivity, design freedom, and high bin-packing density [1,2,3,4]. However, BJT is a multi-step AM technology requiring subsequent processes such as debinding and sintering or infiltration for a densified final product [5]. In addition, particle sizes larger than those in the metal injection molding (MIM) process are used more generally [1,6,7,8]. The internal pores are well developed because there is layer upon layer of material [9]. Although this feature is advantageous in the debinding process, it can be a disadvantage in the sintering process [9,10,11]. Contrarily, well-developed pores provide advantages for the molten metal infiltration process. In particular, the molten metal infiltration process [12,13] uses the wettability and capillary force between different materials. This process can provide functionality by using the characteristics of different materials depending on the purpose.

Moreover, another issue that must be resolved for the commercialization of BJT AM is cost [14]. One of the reasons is that the BJT AM process’s components are not as efficient as those of existing market products in terms of quality, productivity, and economics [15,16,17,18]. In particular, the gas-atomized powder used as a raw material has a narrow particle size range of 30 to 60 μm, which increases the unit price. Therefore, efficient cost savings can be achieved by using powder particles with a wide particle size range.

In this study, attempts were made to determine the maximum range of powders applicable to the binder jetting process and to optimize post-AM process conditions with high mechanical properties. In addition, using these results and considering the characteristics of the BJT AM process mentioned above, journal bearing parts with high lubrication characteristics, which are difficult to implement with existing processing processes, were manufactured. Existing commercialized parts use liquid lubricants added to cylindrical Fe-C materials, but due to the nature of the powder process, uniform pore distribution is difficult, so it is difficult to distribute the lubricant uniformly. However, the AM process has a high degree of freedom of shape and can create shapes that are difficult to process, so it can be modeled to enable uniform lubricant application inside the journal bearing. Additionally, compared to existing Fe-C materials, Fe-Cu-C materials are known to be self-lubricating and exhibit high mechanical properties [19,20]. Binder jetting uses a powder bed, so the use of different materials is limited. However, because the well-developed pores and the melting points of the materials are different, Fe-Cu densified parts can be manufactured using the Cu infiltration process. In addition, due to the nature of the binder used, a small amount of residual carbon is distributed throughout the material, making it possible to manufacture parts with Fe-Cu-C components. Additionally, to evaluate the possibility of expanding the application field of the BJT-AM process and the infiltration process, a rig test was conducted at an actual excavator company using a Cu-infiltrated journal bearing with a changed design.

2. Materials and Methods

The powder material used in this study was 316L stainless steel powder with a wide application range, and the properties of the powder are shown in Table 1. Powder with a particle size of 125 μm or less by gas atomization and powder with a particle size of 15 to 106 μm, excluding fine powder and large-particle-size powder obtained through sieving, were used. The liquid binder used in this work was ExOne’s Aquafuse.

Figure 1a,b show specimens for evaluating wear characteristics and mechanical strength after binder jetting and Cu infiltration. Figure 1a shows the specimen’s shape for the ball-on-flat wear test, and Figure 1b shows the specimen model for the tensile test, which was processed into a standard specimen and evaluated.

Figure 1c shows the geometries of various patterns formed inside a journal bearing to determine the optimal internal geometry through rig testing. Nine types of journal bearings were designed by placing internal circular protrusions inside the journal bearings to allow the lubricant to easily penetrate and by varying the diameter and presence or absence of internal dimples.

BJT AM was performed using M-Flex equipment (ExONE, North Huntingdon, PA, USA), and the process variables used are shown in Table 2. After the process was completed, the powder was fully cured in an oven at 473 K for 8 h, and then the powder was removed.

The molten Cu infiltration process was tested by applying the Cu content of the maximum infiltration weight, which was calculated based on the porosity of each sample pre-sintered at different temperatures. Cu powder was uniaxially compressed, placed on the pre-sintered part, and fed into a continuous heating furnace. Pre-sintering was performed at 1373 K, 1423 K, 1473 K, 1523 K, and 1573 K, and Cu infiltration was carried out at 1373 K for 30 min under a nitrogen–hydrogen mixed-gas atmosphere. In subsequent testing and characterization, a polishing process was added to remove any Cu infiltrate residue that may remain on the surface or to ensure the accuracy of the evaluation even in the absence of it.

A field emission scanning electron microscopy (FE-SEM, FEI Quanta200F, Waltham, MA, USA) analysis was performed to observe the feedstock powder and microstructure images of each specimen. Phase composition was investigated by X-ray diffractometry (XRD, SmartLab, Rigaku, Japan). The particle size of the powder was measured using the LASER particle size analysis (Mastersize 2000, Malvern, UK) method. All specimens were measured for density using the Archimedean method. The Vickers hardness (HMV-2 TE, Shimadzu, Japan) and tensile strength (AG-300kNX Plus, Shimadzu, Japan) of the specimens were measured. The pin-on-disk (ASTM G99) method and ball-on flat (ASTM G133) method were used for wear tests. Lastly, the journal bearing, whose internal design was changed, underwent a rig test by construction equipment manufacturer Doosan Infracore Co., Ltd., and an internal shape that provides optimal load-carrying performance and wear resistance characteristics when installed on an excavator was selected.

3. Results and Discussion

Figure 2 shows the particle size distribution, particle morphology, and XRD analysis results of the powder. The delta ferrite phase is observed in the XRD measurement results due to the atomization process of high-temperature melting and rapid cooling. As the Cr-Ni equivalent in the composition decreases, delta ferrite is formed, and the formed ferrite reverts to austenite due to the rapid cooling rate. It is believed that austenite and ferrite coexist even after cooling to room temperature due to failure to transform. These results are based on a study by Laxminarayana et al. [21], who studied the formation of delta ferrite in austenitic stainless steel casts, including 316L and 304, and based on gas atomization and plasma spheronization treatment to produce 304L spherical powder for the SLM process. This is in good agreement with the study by Sehhat et al. [22]. According to these authors, the melting process can change into austenite–ferrite solidification mode or ferrite–austenite solidification mode depending on the change in chromium–nickel equivalency, and the generated ferrite cannot be transformed back into austenite under a fast cooling rate, so it can be stored at room temperature. It is reported that even after cooling, austenite and ferrite coexist.

Figure 3 displays a heat map based on measured density by location on the powder bed. Figure 3a shows that the bed density is not uniform when using full-range powder (particle size range of 125 μm or less) manufactured through gas atomization. In contrast, Figure 3b shows that powders with particle sizes of 15–106 μm exhibit a uniform powder bed density. This indirectly shows that a stable process is possible even with cheaper powder (with a wide particle size range) than the powder used in the existing BJT AM process.

To optimize Cu infiltration and pre-sintering conditions, the change in the shrinkage rate with the pre-sintering temperature and the microstructure before/after infiltration were compared, and the results are shown in Figure 4 and Figure 5, respectively. As shown in Figure 4, as the pre-sintering temperature increases, the length change in each axis increases due to increased shrinkage in the part. Specimens pre-sintered at 1373 K show isotropic shrinkage, with shrinkage being within 2% regardless of the axis. This is because at low temperatures, there is little change in the position between the powders, and slight necking begins at the contact surfaces, which does not affect the overall part’s dimensions. However, as the temperature before sintering increases, the amount of shrinkage increases, especially in the case of the Z axis, where it can be seen that the amount of shrinkage increases further because it is also affected by gravity. In particular, when pre-sintering is performed at temperatures above 1523 K, very large Z axis shrinkage can be seen to occur irregularly. Table 3 summarizes the density measurement results after pre-sintering measured using the Archimedes method. In the infiltration process, the infiltration amount was calculated and tested based on the density of the pre-sintered part.

Figure 5 shows macroscopic images and the microstructure results before and after Cu infiltration according to the pre-sintering conditions. Figure 5a macroscopically shows the shape of the sample before and after the infiltration process. After infiltration, the front depth of the infiltrant shows 50%, 75%, and 100% infiltration in proportion to the weight of the filler. The left and right sides of the microstructure show images before and after infiltration, respectively. The Cu infiltration results on the right side of each figure were obtained in BSE mode. The bright part represents Cu. The relatively dark part is the SS316L laminated part. Figure 5b shows the microstructure of a specimen pre-sintered at 1373 K. It can be seen that because the sintering temperature is low, densification is not achieved, and large pores remain. Infiltrated Cu fills the pore channels by capillary action, but it is considered that a channel that is too large cannot fill all of the empty pores due to a weak capillary force. In addition, as can be seen in the structure on the right, cracks or pore channels inside the sample grow, and porosity increases due to some Cu-filled pore channels and shrinkage behavior during the infiltration process. On the other hand, the decrease in pore size due to shrinkage at temperatures between 1423 K (Figure 5c) and 1523 K (Figure 5e) indicates an increase in capillary force, resulting in an adequate penetration of Cu. However, for the powder pre-sintered at 1573 K (Figure 5f), the high sintering temperature resulted in the formation of some closed pores, resulting in partially incomplete Cu infiltration.

Table 3 shows the density before/after Cu infiltration according to the pre-sintering temperature. Similar to the microstructure results, the sintered density increases as the pre-sintering temperature increases, and after Cu infiltration, the density increases compared to that of the sample pre-sintered regardless of the pre-sintering temperature. Nevertheless, the density of the sample with Cu infiltration after pre-sintering at 1573 K is higher than that of the 1523 K pre-sintered body with high Cu infiltration. This is believed to be due to the densification behavior with the spherification of pores and grain growth due to high-temperature sintering, as well as the effect of Cu infiltration through the remaining pore channels.

Figure 6a shows the Vickers hardness measurement results according to the Cu infiltration results. The hardness value represents the average value obtained from randomly measuring the hardness at five locations. The samples pre-sintered at 1373 K and 1573 K showed poor Cu infiltration (Figure 5a,e) and lower hardness compared to the Cu-infiltrated sample without pre-sintering. And the hardness of the Cu-infiltrated specimen after pre-sintering at 1473 K was the highest, 230 HV. Figure 6b shows the tensile strength measurement results of the Cu-infiltrated samples according to the pre-sintering temperature. The tensile strength tends to increase as the temperature increases because the higher the pre-sintering temperature, the higher the sintered density (Table 3). The specimens Cu-infiltrated after pre-sintering below 1423 K showed relatively low tensile strength. This is believed to be due to insufficient Cu infiltration and low sintering density, as explained previously.

The results in Table 3 and Figure 6 show that the density and tensile strength after Cu infiltration were highest in the sample pre-sintered at 1573 K; however, the hardness value was higher in the samples pre-sintered from 1423 K to 1523 K with a relatively large amount of Cu infiltration but relatively low density. As is well known, when Cu is infiltrated into an iron-based alloy, the amount of infiltrant cannot exceed the open pore volume, but it is known that as the amount of infiltrated Cu increases, the mechanical properties, including hardness, improve [23,24,25]. In addition, Fe-Cu-C, which can be formed by the reaction of infiltrated Cu and residual carbon, is an alloy composition used for high load-bearing parts due to its high hardness and wear resistance. Therefore, the hardness tends to decrease even as the density increases, which is presumed to be caused by a combination of the Cu infiltration effect and the formation effect of Fe-Cu-C.

Ball-on-flat wear tests were performed to observe the effect of Cu infiltration on wear resistance, which is one of the most important objectives. Measurements were performed separately for the pre-sintered and Cu-infiltrated samples after pre-sintering. For the pre-sintered sample (Figure 7a), the friction coefficient increased with an increasing pre-sintering temperature. This is believed to be because the density increases with an increase in the sintering temperature. However, in the case of the Cu-infiltrated sample (Figure 7b), even though no lubricant was used, a lower coefficient of friction was obtained in all samples compared to the whole pre-sintered sample regardless of the sintering temperature. Therefore, this improvement in wear resistance is believed to be due to the formation of Fe-Cu-C with self-lubricating properties through Cu infiltration [19,20], as described in the Introduction.

As a result of combining the results of each characteristic, the conditions of pre-sintering at 1473 K and then Cu infiltration were selected as post-process conditions suitable for manufacturing journal bearings. Figure 8 shows nine journal bearing parts manufactured through the BJT AM process based on the model in Figure 1 and then pre-sintered at 1473 K and infiltrated with Cu. The excavator rig testing results show that the journal bearings manufactured by BJT AM exhibited superior wear resistance and load-bearing capacity compared to commercial products. The specimen that showed the best mechanical properties was specimen #9, which had an internal dimple on the internal pattern protrusion with an inner diameter of 3 mm and an outer diameter of 5.6 mm. In particular, the developed bearings were found to have higher wear resistance and load-bearing capacity compared to existing commercial products. The wear amount and load-bearing capacity of the developed product (#9) were 0.11 mm and 117.7 N/mm², respectively, which were both superior to the 0.65 mm and 68.6 N/mm² values of the commercial product.

4. Conclusions

In this study, journal bearing products for excavators were studied to expand the scope of application of the BJT AM process. In particular, the costs of the process can be reduced by applying a wide range of raw material powder particle sizes. In addition, we developed journal bearing parts with improved mechanical properties through patterning utilizing the BJT AM process’s characteristics and conducted rig tests for commercialization. The conclusions are as follows:

• The maximum particle size range applicable to the BJT AM process using 316L powder manufactured by gas atomization was determined to be 15–106 μm.

• A post-AM design with the most advantageous mechanical properties in the 316L Cu composite was studied, and the pre-sintering temperature was determined to be 1473 K.

• The actual excavator heavy equipment rig testing results show that the mechanical properties of products with internal patterns formed through BJT AM improved by more than 50% compared to those of mass products without internal patterns.

Data Availability Statement

The original contributions presented in the study are included in the article, further inquiries can be directed to the corresponding author.

Conflicts of Interest

The authors declare no conflicts of interest.

Footnotes

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Figures and Tables

Figure 1. Test artifact standard model. (a) Ball-on-flat test specimen, (b) tensile strength test specimen, and (c) 3D models of nine types of journal bearing.

Figure 2. Feedstock powder properties. (a) Powder morphologies and (b) XRD results.

Figure 3. Heat map image of powder bed density survey: (a) size under 125 μm; (b) powder size in the range of 15–106 μm.

Figure 4. Length deviation of each axis after pre-sintering processes: (a) deviation of 1373 K pre-sintered part to green body part, (b) deviation of 1423 K pre-sintered part to green body part, (c) deviation of 1473 K pre-sintered part to green body part, (d) deviation of 1523 K pre-sintered part to green body part, and (e) deviation of 1573 K pre-sintered part to green body part.

Figure 5. Macroscopic images and microstructure images before/after molten Cu infiltration process (BSE mode): (a) macroscopic images before/after infiltration, (b) pre-sintering part at 1373 K/Cu infiltration part, (c) pre-sintering part at 1423 K/Cu infiltration part, (d) pre-sintering part at 1473 K/Cu infiltration part, (e) pre-sintering part at 1523 K/Cu infiltration part, and (f) pre-sintering part at 1573 K/Cu infiltration part.

Figure 6. Mechanical properties of pre-sintered and Cu-infiltrated specimens. (a) Vickers hardness test. (b) Tensile strength.

Figure 7. Wear testing under all process conditions: (a) ball-on-flat friction coefficient measurement result for each pre-sintering process condition and (b) ball-on-flat friction coefficient measurement result for molten Cu infiltration after pre-sintering.

Figure 8. Nine journal bearings manufactured using the BJT AM process.

Table 1

Powder characterization.

Particle Size Range (μm)	d10	d50	d90	Apparent Density (g/cm3)	Tap Density (g/cm3)	Flowability (Hall, sec/50 g)
<125 μm	11.4	28.4	92.7	4.86	5.41	-
15~106 μm	18.9	36.3	78.8	4.29	5.30	20.97

Table 2

Process parameters.

Dispensing Speed (mm/s)	Spreading Speed (mm/s)	Binder Saturation Ratio (%)	Bed Temperature (K)	Drying Speed (mm/s)
150	5	60	328	10

Table 3

The density of each specimen was measured by the Archimedes method.

Pre-Sintering Temperature (K)	Relative Density (%)
	Pre-Sintered	Cu Infiltrated
Green body (not sintered)	-	88.31 ± 1.68
1373	71.75 ± 1.60	90.68 ± 1.16
1423	73.93 ± 2.00	91.45 ± 0.85
1473	76.38 ± 1.41	92.04 ± 0.82
1523	77.71 ± 1.92	93.72 ± 1.03
1573	80.69 ± 1.36	95.69 ± 0.80

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