Erosion-Wear Experimental Materials and Equipment

Experimental Materials

Tungsten carbide (WC) carbide?is a composite material produced using powder metallurgy techniques, with WC, a metal carbide that is difficult to dissolve, as the matrix and a binder added. It is characterized by high hardness and strong wear resistance. The WC carbide?used in this experiment is YG8, which is employed as the coating material for the valve core and outlet sleeve in coal direct liquefaction devices. YG8 is a tungsten-cobalt carbide?with a cobalt binder, a density of 14.6 g/cm3, a hardness of HV 1350, an elastic modulus of 540 MPa, a bending strength of 1500 MPa, and a compressive strength of 4470 MPa.

The experiment uses quartz sand (SiO?) particles as the erosion particles, which are commonly used in erosion-wear tests. The particles are sieved to achieve an average particle size of 150 μm, as shown in Figure 1.

Experimental Equipment

The experiment uses a self-made shock wave-driven gas-solid two-phase flow erosion-wear testing device. This device primarily consists of a shock wave generator, a velocity measurement system, a high-speed camera, a heating system, and a temperature control system.

In this setup, nitrogen gas is connected to the driving section to generate shock waves with a specific Mach number. The driving section and the driven section are separated by an aluminum film. Experimental particles are placed on a tin foil located between the driven section and the accelerating section. By adjusting the pressure relief valve, gas is introduced into the driving section of the shock tube. When the pressure difference across the film reaches a critical value, the film ruptures suddenly, generating a shock wave. The high-speed gas flow then propels the solid particles through the accelerating section to the desired experimental velocity. Upon impacting the specimen surface, the high-speed particles cause material loss, thus facilitating the erosion process.

The driving section, driven section, and accelerating section are each equipped with dynamic pressure sensors, charge amplifiers, and dynamic test analyzers to measure shock wave velocity. A high-speed camera in the experimental section captures the particle motion trajectories. The specimen holder is equipped with a temperature heating system and a temperature control system, allowing for adjustment to the required experimental temperature.

Erosion-Wear Experimental Parameters and Methods

Experimental Parameters

The impact angle, denoted as θ (see Figure 3), is defined as the angle between the axis of the shock tube and the surface of the specimen being eroded (0°to 90°). The desired impact angle is achieved by rotating the specimen holder.

The impact distance L is the distance between the center of the shock tube’s outlet and the center of the specimen’s surface. Typically, L is set between 30 and 50 mm during experiments. When conducting experiments at different impact angles, the position of the shock tube needs to be adjusted to maintain a consistent impact distance.

In the experiment, the impact velocity is adjusted by changing the thickness of the aluminum foil. Aluminum foils with thicknesses of 0.13 mm, 0.20 mm, and 0.30 mm are used for the erosion-wear tests. A high-speed camera is employed to record the particle trajectories. By analyzing these trajectories, the particle velocity v p is determined.

In the equation, ΔI represents the distance between the ends of the particle clusters in consecutive frames, measured in meters; Δn is the number of frames between measurements; and f is the filming frequency, measured in frames per second (FPS).

Using the high-speed camera, the velocity of 150 μm SiO? particles is tested. By replacing aluminum foils of different thicknesses, the corresponding membrane rupture pressure ratios are obtained, which in turn allows for the determination of the impact velocity. The velocities corresponding to different aluminum foil thicknesses are summarized in Table 1. The specific calculation method for particle velocity can be found in the referenced literature.

Experimental Method

Before the experiment, the nickel-based carbide specimens are first polished using 1000# sandpaper. The specimens are then cleaned with an ultrasonic cleaner, air-dried, and weighed to obtain an average value. The specimens are fixed onto the specimen holder, and the angle of the specimen and the distance between the shock tube and the specimen are adjusted. The temperature control system is activated, and the experimental temperature is set. Aluminum foils of the appropriate thickness are selected and solid particles are loaded simultaneously.

The lighting is turned on, and the dynamic testing analyzer and high-speed camera are activated. The camera lens height is adjusted so that the distance from the shock tube outlet to the specimen surface is within the field of view of the high-speed camera. The nitrogen gas valve is then opened to start the experiment. When the aluminum foil ruptures, the valve is immediately closed, and the high-speed camera captures the particle trajectories during the experiment.

At the end of the experiment, the specimen is cooled, cleaned, and dried. The specimen is weighed 10 times using an electronic balance to record the average weight. The erosion-wear rate is then calculated using equation (2).

In the equation, E represents the erosion-wear rate, measured in mg/g; Δm is the mass loss of the material, measured in mg; and m p is the mass of a single impact particle, measured in grams.

Analysis and Discussion of Erosion-Wear Experimental Results

Effect of Impact Angle on YG8 Wear Rate

Under an impact velocity of 175 m/s, the erosion-wear rates of the specimens were measured by varying the impact angles, as shown in Figure 4.?

From Figure 4, it can be observed that the erosion-wear rate of the specimen initially increases and then decreases as the impact angle increases. The erosion-wear rate reaches its peak at an impact angle of 75°. The experimental results indicate that YG8 is a typical brittle material, with the maximum erosion-wear rate occurring at high impact angles. The erosion-wear characteristics of YG8 are consistent with the behavior of brittle materials, where the erosion-wear rate varies with the impact angle.

Impact Angle on Coating Erosion-Wear Performance

Particle velocity is a crucial factor affecting the wear rate of materials. Impact experiments were conducted on specimens at impact angles of 30°, 60°, and 90° under three different impact velocities: 148 m/s, 175 m/s, and 200 m/s. The relationship between erosion-wear rate and particle impact velocity is shown.

Figure 5 demonstrates that, at all three impact angles, the erosion-wear rate of the material increases with increasing impact velocity. There is a critical impact velocity at which erosion-wear begins, related to the abrasive properties and the material’s characteristics. Erosion-wear occurs only when the velocity exceeds this critical value. Extensive erosion tests indicate that the erosion-wear rate has the following relationship with particle velocity:

E=kv” （3）

where

v is the particle velocity in m/s;

k is a constant; and

n is the velocity index. A higher

n value indicates that the erosion-wear rate of the material is more influenced by the particle impact velocity.

Fitting the experimental data to equation (3) yields velocity indices of 2.34, 2.27, and 2.28 for impact angles of 30°, 60°, and 90°, respectively, for YG8 material.

Analysis of Erosion-Wear Mechanisms

Analysis of the erosion-wear morphology of specimens at an impact angle of 90° reveals that the erosion-wear is primarily driven by the impact forces of the solid particles directly striking the composite layer. Due to the high brittleness of both tungsten carbide particles and the composite layer matrix, high-velocity solid particle impacts more readily induce plastic deformation or crack formation, leading to the development of pits and cracks.

During the erosion process, the abrasive quartz sand continuously impacts the surface, creating numerous pits. The edges of these pits accumulate material that has been deformed and squeezed out, forming a lip-like flange. With continued particle impacts, this flange is progressively eroded and stripped away due to repeated compression. The wear mechanism can be summarized as erosion-induced compression leading to pit formation and material detachment.

As the reinforcement phase, WC particles have much higher hardness and stiffness compared to quartz sand, which helps them better withstand the abrasive impacts. During the erosion-wear process, the coating undergoes cutting and plowing effects from the sharp edges of the abrasive particles, resulting in plastic deformation, progressive fatigue, and delamination. The protruding WC particles bear the brunt of the abrasive impact. Table 2 shows the elemental chemical composition of the specimen surface before and after the erosion experiments.

Conclusion

As the impact angle increases, the erosion-wear rate of YG8 material first increases and then decreases, reaching a maximum at an impact angle of 75°. YG8 exhibits the erosion-wear characteristics typical of brittle materials.

At impact angles of 30°, 60°, and 90°, the erosion-wear rate of YG8 material increases with rising impact velocity. The corresponding velocity indices, obtained from fitting the erosion-wear rate versus velocity relationship, are 2.34, 2.27, and 2.28, respectively.

The primary erosion-wear mechanism for YG8 material involves the formation of pits and microcracks on the material surface due to high-angle impacts. These features are caused by the detachment of Co and WC particles from the matrix and the development of microcracks under high-velocity impacts.

Requirements for Seal Face Materials in Mechanical Seals

To ensure the proper operation of sealing rings in mechanical seal devices, they are typically configured as a pair consisting of a hard ring and a soft ring with different hardness levels, considering aspects such as wear reduction, corrosion resistance, and prevention of galling. During operation, the sealing rings may come into contact and generate friction when starting, stopping, or experiencing fluctuations in working conditions. Therefore, the material for the hard ring needs to have sufficient strength, rigidity, wear resistance, and thermal conductivity. Friction and fluid shear forces can elevate the temperature of the sealing rings, so the sealing material must exhibit good thermal conductivity, heat resistance, and thermal shock resistance. To ensure a long service life, the sealing ring must also have good corrosion resistance. Additionally, the hard ring should possess good formability and machinability, low density and permeability, and excellent self-lubricating properties. No single material can fully meet all these requirements, so typically, the main performance criteria for sealing materials are defined based on the operating environment, and suitable materials are selected accordingly.

Application of WC-Ni carbide as Mechanical Seal Materials

To ensure the longevity and stable operation of mechanical seal materials, the sliding materials should have appropriate thermal compatibility and thermal conductivity, as well as suitable coefficients of thermal expansion, elastic modulus, and friction factors. WC-Ni carbides are known for their excellent performance in mechanical seals, making them suitable for applications in high-pressure, high-speed, high-temperature, corrosive environments, and media containing solid particles.

Characteristics of WC-Ni carbide

Since the introduction of carbides in the 1920s, cobalt has been regarded as the best binder phase and continues to play a significant role in the preparation of carbides. With the rapid advancement of science and technology, the applications of carbides have expanded, leading to a surge in demand. Due to the scarcity of cobalt resources, scientists worldwide have prioritized cobalt as a strategic material and have been researching ways to reduce or substitute cobalt in carbides. Nickel, being close to cobalt in the periodic table, with similar density, melting point, and atomic radius, can effectively wet and support the hard phase and has lower radioactivity compared to cobalt, making it a common substitute.

The characteristics of WC-Ni and WC-Co carbides during the sintering process are similar. However, due to the different strengthening effects of Ni and Co on the hard phase, WC-Ni may exhibit slightly lower performance in certain aspects compared to WC-Co carbides. By adding a small amount of metal elements to enhance the binder phase, using fine low-carbon WC particles, and employing vacuum sintering processes, WC-Ni carbides with lower porosity and a uniform, fine-grained structure can be achieved. Their hardness, bending strength, and tribological properties can meet or exceed those of WC-Co carbides, while their corrosion resistance is also significantly improved. Additionally, as Ni replaces the radioactive element Co, it provides good radiation protection when used under radioactive conditions. By carefully controlling the total carbon content and grain size of the alloy, and adding appropriate amounts of Mo and Cr, WC-Ni carbides can be produced with non-magnetic properties and excellent physical and mechanical performance, thereby mitigating the effects of special working conditions and environmental factors.

Physical and Mechanical Properties of WC-Ni carbide for Mechanical Seals

WC-Ni carbides are made by mixing WC and Ni powders in a specific ratio, adding a binder, and then pressing and sintering the mixture. With a melting point of approximately 2700°C, WC particles are primarily bonded together during the sintering process through the melting of Ni. At high temperatures, some WC dissolves into Ni, forming a WC-Ni eutectic with a lower melting point than Ni. Consequently, the sintering temperature varies with changes in Ni content and WC grain size. For composite materials, physical parameters such as elastic modulus, coefficient of thermal expansion, Poisson’s ratio, thermal diffusivity, and thermal conductivity can vary based on the proportion and distribution of each phase.

Application of WC-Ni carbide Materials in Mechanical Seals

Due to their exceptional toughness, rigidity, high hardness, good wear resistance, high bending strength, and high thermal conductivity, both WC-Ni and WC-Co carbides are notable. WC-Ni carbides offer superior corrosion resistance compared to WC-Co alloys and do not emit radiation under neutron exposure, making them suitable for use in mechanical seals operating under high pressure, high speed, high temperature, corrosive media, media containing solid particles, and radioactive environments. Currently, WC-Ni carbides have significant application value in vehicle transmission shaft seals, power shift transmissions, pumps in special operating conditions, and rotary seals for aircraft, as well as in the petrochemical industry and nuclear power seals.

Influence of Microstructure on the Properties of WC-Ni carbides

The unevenness in microstructure can adversely affect the strength of carbides. Minor variations in the binder phase content and distribution, WC grain size, carbon content, and any form of impurity contamination can lead to an uneven microstructure that negatively impacts the mechanical properties of WC-Ni carbides.

Influence of WC Grain Size on the Properties of WC-Ni carbides

WC-Ni carbides use Ni as the binder metal. During the sintering process, Ni melts at the sintering temperature and bonds the WC particles together into a solid mass. These alloys exhibit very high hardness, are difficult to machine, and possess excellent wear resistance. Variations in the processing methods can lead to significant differences in the alloy’s composition and properties, and the morphology of the WC grains can also affect the performance of WC-Ni carbides.

3.2Changes in WC Grain Morphology During Liquid Phase Sintering of WC-Ni carbides

The coarseness of WC grains can significantly affect the bending strength of carbides, while uneven distribution of Ni can lead to brittle fracture of the alloy. To improve the fracture toughness of the product, it is essential to strengthen the interface between WC and the binder phase or to enhance the strength of the binder phase. Therefore, controlling the sintering process and conditions will impact the mechanical properties of WC-Ni carbides.

During sintering, the shape of WC grains in the carbide?is also influenced by shape relaxation and the grain growth process. The higher the ratio of the average intercept length of the binder phase (the average length of each grain intersected by any testing line on a cross-section) to the WC grain size, the less impact it has on the shape of the WC grains, resulting in a more equiaxed grain morphology.

Effect of Binder on Temperature Residual Stress in WC-Ni carbides

When the content of the metallic binder Ni in WC-Ni alloy materials is relatively high, the compressive stress in fine WC grains is greater than that in coarse grains. This is because, with a constant WC content, the average free path of the binder in fine powder is shorter than in coarse powder. When the WC-Ni alloy contains less binder, the difference in the average free path of the binder is minimal, and the variation in residual stress with temperature is not significant. Therefore, if conditions allow, the Ni content in WC-Ni alloy sealing rings should be reduced to minimize the uneven distribution of residual stress due to temperature changes, thus reducing or even preventing thermal cracking in the sealing rings.

Current Research Status on the Application Performance of WC-Ni carbides

?Corrosion Resistance of WC-Ni Alloys

Compared to WC-Co carbides, WC-Ni carbides exhibit superior wear resistance. This is due to the binder in WC-Ni alloys having excellent corrosion resistance, with both passivation and electrochemical corrosion rates for WC-Ni carbides being significantly lower than those for WC-Co carbides. Under acidic conditions in practical production processes, WC alloys with Ni as the binder show better acid resistance than those with Co as the binder.

Table 2 compares the corrosion resistance of WC-Ni and WC-Co carbides. The results show that substituting Co with Ni significantly enhances the corrosion resistance of WC carbides. However, the corrosion resistance of a material is specific to its alloy composition, grain size, and the corrosive conditions (including temperature, concentration, time, and corrosion state). For example, the corrosion resistance of YWN8 in 68%-90% HNO? is not significantly different from, and even slightly lower than, that of YG6 alloy. This is primarily due to the poor resistance of metallic Ni to strong oxidizing acids like HNO?; as the concentration of HNO? and the Ni content in the alloy increase, its corrosion resistance decreases.

Table 2 Corosion resistance performance comparison between WC-Ni and WC-Co cemented carbides

Tribological Performance of WC-Ni Alloy Mechanical Seal Rings

WC-Ni carbide?sealing rings exhibit excellent wear resistance. This is because WC-Ni carbides possess strong oxidation resistance and corrosion resistance in fluid sealing media, which contributes to their superior wear resistance. The friction coefficient of WC-Ni carbides is related to the content, grain size, and distribution of the binder phase. A softer binder phase can lead to adhesion during friction. Additionally, the content and composition of the binder phase can affect the hardness of WC-Ni, thereby influencing the wear resistance of the WC-Ni carbide.

Conclusion

Mechanical seal materials are a crucial area of research in sealing technology. With the advancement of modern science and technology and the increasing demands of production and daily life, the requirements for sealing technology have become more stringent. However, research in this field in our country is still relatively behind. Strengthening interdisciplinary collaboration and continually improving experimental and theoretical research are key to overcoming the technological barriers in mechanical seals imposed by foreign countries.

Experimental Methods

Raw Materials

The experiment uses coarse and extra-coarse WC powders from well-known suppliers, with their main characteristics shown in Table 1. Additionally, 2.0 μm cobalt powder from the same supplier was also used.

Experimental Methods

For the preparation of WC-10%Co (where all content is given in weight percentage), weigh 900 g of WC, 100 g of Co, and 20 g of PEG. Measure 235 mL of alcohol and 2000 g of grinding balls. Add these into a 2.4 L ball mill. The mill is operated at a speed of 63 r/min for 14.5 hours. After milling, the mixture is dried, sieved, and then pressed into samples weighing 10 g each. The samples are sintered in a continuous vacuum sintering furnace at 1450°C.

Particle Size Measurement

For coarse tungsten carbide, measure the Fischer particle size in both the as-supplied and milled states. The samples are resin-mounted and analyzed using a metallurgical microscope to determine the grain size and particle size distribution of the powder. The alloy grain size and particle size distribution are measured using classic metallographic methods, and the coercive force of the samples is also assessed.

Results and Analysis

Fischer Particle Size (Fsss) and Alloy Grain Size

As-Supplied Particle Size and Alloy Grain Size

The metallographic images of alloys made from WC powders #1 and #2 are shown in Figures 1 and 2, respectively. Comparing Figures 1 and 2, it can be observed that the WC grain size in Figure 2 appears to be slightly coarser than in Figure 1. This indicates that coarser as-supplied Fsss particle sizes of WC lead to coarser grain sizes in the WC-Co alloys. Metallographic analysis shows that the average WC grain sizes for alloys made from powders #1 and #2 are 4.8 μm and 5.8 μm, respectively. Thus, the average grain size of WC in sample #2 is 1.2 times that in sample #1. The as-supplied Fsss particle size of #2 WC powder is 2.5 times that of #1 WC powder. Clearly, there is no direct proportional relationship between the as-supplied Fsss particle size of WC powder and the alloy grain size. Additionally, the Fsss particle size values for #1 WC powder are 2.5 times the alloy grain size, and for #2 WC powder, it is 5.3 times the alloy grain size. This indicates that the as-supplied WC powders for both samples are primarily aggregated polycrystalline WC particles, with more severe agglomeration for coarser WC powders.

Milled Particle Size and Alloy Grain Size

Comparing Table 1 with the metallographic grain sizes in Figures 1 and 2, it can be seen that the Fsss particle sizes of milled WC powders #1 and #2 are relatively close to the alloy grain sizes. Moreover, the measured alloy grain sizes are higher than the milled Fsss particle size values. This discrepancy is due to differences in measurement principles as well as grain growth during the sintering process. However, it clearly indicates that the Fsss particle sizes of coarse WC powders in the milled state are very close to the alloy grain sizes. The ratios of average grain sizes to milled particle sizes for alloys #1 and #2 are 1.15 and 1.31, respectively.

Raw Material WC Grain Size and Alloy Grain Size

Results from Direct Metallographic Measurement

Metallographic images of #1 and #2 WC powders after mounting and etching are shown in Figures 3 and 4. The grain sizes measured using metallographic methods are 5.31 μm and 8.5 μm, respectively. The grain size distributions of the powders and alloys are shown in Figures 5 and 6.

Figures 3 and 4 clearly indicate that the grain size of #2 WC is significantly larger than that of #1 WC. This suggests that WC with a coarser as-supplied Fischer particle size also has coarser grains. Additionally, it is evident that #1 WC exhibits better dispersion, with less pronounced sintering between particles compared to #2 WC. The severe sintering in #2 WC particles is a major reason why the metallographic grain size is much larger than the alloy grain size, and also explains why the grains in #2 WC are much larger than those in #1 WC.

From the grain size distribution of the raw powders and alloys in Figures 5 and 6, it can be seen that sample #1 contains coarse WC grains of 15–20 μm in the raw material, which are not present in the alloy. In contrast, sample #2 has a substantial amount of WC grains in the 15–35 μm range, though only a small amount of 15–20 μm grains are found in the alloy. This suggests that the severe sintering of the mounted WC, although difficult to distinguish by metallographic methods after etching, was fragmented during the intense grinding process.

Moreover, comparing the WC and alloy grain size distributions in Figures 5 and 6 shows that the grain size distribution of WC in sample #1 is more consistent with the alloy grain distribution than in sample #2. This consistency is a significant reason why many researchers believe that WC similar to sample #1 is more conducive to producing coarse alloys with a more uniform grain size.

WC Particle Size and Alloy Coercive Force

The coercive forces of the alloys made from #1 and #2 powders are 4.6 kA/m and 4.3 kA/m, respectively. The relationship between the WC-Co alloy grain size and the alloy’s coercive force can be expressed using the empirical formula (1).

In the formula:

• Hc= coercive force of the alloy (kA/m)
• Com= cobalt content in the alloy (%)
• Dwc= average WC grain size in the alloy (μm)

According to the calculations, the average grain sizes of alloys #1 and #2 are 7.4 μm and 8.8 μm, respectively. Clearly, the calculated grain sizes are significantly larger than the measured grain sizes, but the difference between the average grain sizes of alloys #2 and #1 is close to the difference observed using metallographic methods. The results obtained from formula (1) do not show a clear quantitative relationship with the Fsss particle sizes of the raw WC in both states, but the size of the raw material particles can still be used to predict the alloy grain size and coercive force.

Conclusions

Based on the above, the following conclusions can be drawn:

1.Coarse WC powders with larger as-supplied Fsss particle sizes tend to have higher milled Fsss particle sizes and larger grain sizes, leading to alloys with larger grain sizes.

2.The Fsss particle size in the milled state of coarse WC can be used to evaluate the grain size of coarse WC and predict the grain size of WC-Co alloys. Under the test conditions, the alloy grain size is 1.1 to 1.3 times the Fsss particle size of the milled WC.

3.Coarse WC powders with as-supplied Fsss particle sizes around 10 μm have a better consistency in grain size distribution with the alloy WC grain size distribution compared to extremely coarse WC powders with Fsss particle sizes above 25 μm.

Generally, compared to traditional insert end mills, solid carbide end mills offer significantly higher machining precision. This fact holds true specifically for the performance of the tool during cutting operations. Although both types of tools may have the same dimensional precision grade, the rigidity of solid carbide material is much higher than that of a steel tool holder with insert blades. This means that solid carbide end bit mills are less likely to undergo deflection or deformation under cutting forces, leading to higher machining precision.

Base Material of Solid Carbide End Mills

The machining performance of solid carbide end mills largely depends on the type of carbide base material used. The base material is crucial because it must support the cutting edges of the tool, endure significant cutting forces, and prevent any form of tool damage.

To ensure that end mills have sufficient toughness and provide good dynamic resistance, solid carbide end mills typically use micro-grain carbide as the base material (see Figure 2). This base material offers higher hardness and better edge sharpness while maintaining good toughness. However, compared to conventional grain-sized carbide, micro-grain carbide has relatively poorer thermal conductivity (its ability to dissipate heat from the cutting area). This means that heat generated during cutting tends to remain on the tool surface. Therefore, the cutting edges of solid carbide end mills must be able to withstand this cutting heat and manage the contact arc length, which is an important consideration when selecting solid carbide end mills.

Coating and Cutting Edge Preparation of Solid Carbide End Mills

To enhance tool wear resistance and isolate the cutting zone, which generates heat, from the tool base (since heat accumulation in the base material can shorten tool life), solid carbide end mills are typically coated. Additionally, because the cutting edges of solid carbide end bit mills are quite sharp, appropriate adhesion between the tool base and the coating is also crucial (see Figure 3). This is especially important for smaller diameter solid carbide end mills, where the sharpness of the cutting edges is a key factor in the tool’s machining performance.

The Ideal Cutting Edge for Solid Carbide End Mills

The ideal cutting edge of a solid carbide end mill should have as high a hardness as possible while minimizing the risk of chipping. This goal can be partially achieved through proper cutting edge preparation. Generally, different solid carbide end mills may employ various cutting edge preparation methods, edge geometries, and sharpness levels depending on the required machining quality and tool life.

The cutting edge is the intersection line between the tool’s rake face and flank face. By grinding these faces, a sharp cutting edge can be obtained. However, if a PVD coating is deposited directly on a sharp cutting edge, it can create high internal stresses within the coating. These high internal stresses can cause the coating to crack and peel off during cutting, thereby shortening the tool’s life. The quality and effectiveness of the coating depend on its ability to withstand and/or reduce the wear rate during cutting. To ensure that the coating adheres more firmly to the cutting edge and to prevent edge damage, it is necessary to perform edge reinforcement (or blunting) treatment (see Figure 4). In other words, to ensure machining stability and achieve coating functionality, a certain degree of edge sharpness must be sacrificed, which in turn extends tool life.

It can even be said that the importance of cutting edge preparation for solid carbide end bit mills outweighs that of base material type and coating technology. Logically, this has a significant impact on the regrinding of solid carbide end mills. After regrinding, if the cutting edge is not re-blunted to restore its initial condition, the full potential of the tool repair cannot be realized. Therefore, considering the high initial cost of solid carbide end mills, it is crucial for the original tool manufacturer or its qualified service centers to handle tool regrinding services.

Machining Strategies for Solid Carbide End Mills

Solid carbide end mills can be categorized into several major types based on their size and geometry, and further subdivided into many specialized subcategories according to different machining ranges. In various tool application areas, design features such as flute geometry, cutting edge angles, rake and relief angles, and helix angles play crucial roles in differentiating among the types of solid carbide end bit mills. This classification guides the selection of both end mills and machining strategies.

So, which machining strategy is the best choice? This depends on the overall machining goals: is your primary aim to maximize productivity and part output, or to minimize tool costs and simplify tool types? Additionally, it also depends on the workpiece and various related factors: is the tool used for slotting, side milling, or a combination of both?

A final consideration is constraints, such as: what is the potential machining capability of the machine tool? How rigid is the workpiece clamping? These factors might become limiting constraints, preventing the use of more advanced machining strategies or more efficient specialized solid carbide end mills.

The correct choice of solid carbide end bit mill depends on multiple factors, with the most important being the adoption of the correct machining strategy. In practice, many constraints cannot be changed: the machining machine, CAM system, and the material, size, tolerance, and shape of the workpiece are fixed constants. However, within the existing machining system framework, machining results can still be influenced by formulating the right machining strategy and using various methods, and by optimizing cutting conditions through adjustments to feed rate, cutting speed, and cutting depth according to the overall machining goals.

Based on the selected focus and technical strategy, appropriate solid carbide end mills can be chosen. There are two obvious ways to select tools:

1. Performance-Based Selection:Choose specialized end mills with specific applications (such as side milling, slotting, or 3D profiling) to achieve optimal performance.
2. Versatility-Based Selection:Choose general-purpose end bit mills with a wider range of applications but fewer types.

Regardless of the selection method, users need to further narrow down the options within the available varieties and specifications of solid carbide end mills.

Summary

The material, size, tolerance, and shape of the machined parts are given constants. However, within the existing machining system framework, machining results can still be influenced by formulating the correct machining strategy and employing various methods. Additionally, cutting conditions can be optimized by adjusting feed rate, cutting speed, and cutting depth according to the overall machining goals.

Based on the chosen focus and technical strategy, appropriate solid carbide end mills can be selected. There are two clear approaches to tool selection:

1. Performance-Based Selection:Choose specialized end mills designed for specific applications (such as side milling, slotting, or 3D profiling) to achieve the best performance.
2. Versatility-Based Selection:Choose general-purpose end mills that, while fewer in types, offer a wider range of applications.

Regardless of the selection method, users need to further narrow down the choices within the available varieties and specifications of solid carbide end mills.

With the rapid development of China’s automotive industry, especially in the booming new energy vehicle sector, lightweighting has become a core topic. The key to lightweighting is changing traditional materials. Aluminum alloys, with their high strength and low weight, are essential for achieving vehicle lightweighting. The complex geometric shapes of automotive parts and the increasing proportion of die-cast aluminum alloy parts in vehicles have led to a growing demand for CNC machining of die-cast parts.

CNC machining of aluminum alloy automotive parts requires high efficiency, stable continuous production, and decreasing costs, necessitating more detailed control and planning of the entire production process.

Tool Build-Up

Aluminum’s inherent property of low melting point causes it to become “sticky” during cutting. Due to this characteristic and inadequate cooling in actual conditions, the heat generated by friction during the microscopic cutting process cannot be released in time or effectively. This results in aluminum melting and adhering to the tool’s cutting edges and chip flutes. Upon cooling, the aluminum solidifies and sticks to the tool, forming build-up. This phenomenon, commonly referred to in the industry as “easy tool sticking,” leads to tool failure.

Tools are consumables in CNC machining processes and represent a significant cost component. Generally, cutting tools for aluminum alloys need to be sharper, with chip flutes specially polished and coated with aluminum-specific coatings to improve chip removal efficiency. The automotive industry’s push for high efficiency increases feed rates and cutting speeds, which raises the heat generated during cutting and the risk of aluminum melting and sticking to the tool, thereby increasing costs due to tool failure from build-up.

With environmental regulations, minimal quantity lubrication (MQL) is widely used as a cutting fluid alternative in aluminum alloy CNC machining. However, the low melting point of aluminum exacerbates the reduced cooling effect of MQL, further promoting build-up. Tools that fail due to sticking account for about 40% of conventional tool failures. Traditional methods of dealing with build-up, such as tapping or striking, rarely restore tools for reuse. Thus, a new solution is proposed.

Treatment Measures

The new solution involves the following steps:

1. Remove the tool with build-up.
2. Obtain solid NaOH, dilute it with water, and place it in a ceramic container.
3. Once diluted into NaOH solution, immerse the tool with aluminum build-up in the solution. Ensure that the build-up areas are fully submerged and maintain immersion for 2 hours, or adjust based on practical conditions. Table 1 compares traditional and new treatment methods.

Chemical Principles

Using A1Si7Mg material commonly found in automotive parts as an example, where A1 content is approximately 93.5%, Si content is 6.5%, and Mg content is 0.25%. Both Al and Si react with NaOH solution. Immersion in NaOH solution removes the primary A1 component from the tool. The principle involves the reaction between metal and NaOH, producing bubbles (H?), which eventually causes the adhered aluminum to detach.

Chemical reaction equations are as follows:

1. Si reacts with NaOH: Si + 2NaOH + H?O → Na?SiO? + 2H?↑
2. Al reacts with NaOH: 2Al + 2NaOH + 2H?O → 2NaAlO? + 3H?↑

The final result is the removal of aluminum, making the tool reusable.

Experimental Validation for Reducing Build-up

The theoretical method was tested using a tap. Taps are valuable tools in aluminum alloy machining, requiring a longer lifespan and featuring complex geometric shapes. Once aluminum adhesion occurs, physical removal is nearly impossible, making this test more significant and representative.

Due to high heat generated during machining and possible inadequate cooling, aluminum melts instantly and adheres to the flutes rendering the tap unusable due to damaged threads.

The test involved immersing the tap with aluminum build-up in NaOH solution.

The test conclusion: The tap, after complete immersion in NaOH, showed that the build-up had completely detached. Residual aluminum slag was found in the test container. The treated tap was used for further machining, and the workpiece threads met the required specifications. The tap was successfully restored for reuse.

Conclusion

The automotive parts industry, characterized by mass production, requires extensive cutting validation during the initial equipment and tool design phase. Common issues such as build-up during validation due to parameter mismatches, equipment adjustments, and operator experience can significantly increase trial costs and production cycles. This method effectively addresses build-up issues, greatly reducing tool costs and processing time, extending tool life, and substantially lowering production costs.

Conventional end mills have the following drawbacks under high-efficiency cutting conditions:

1.To reduce cutting forces and facilitate chip evacuation, the end edges of corner radius end mills typically feature a concave design with a high edge at the tip and a low center. This means that only the tip participates in cutting during face milling, resulting in high stress and a tendency for chipping.

2.The cutting edge of ball-end mills exhibits both negative rake angles and low-speed cutting inefficiencies, leading to a low metal removal rate.

3.Cutting forces are primarily radial, with the main cutting forces directed along the X and Y axes, causing tool chatter under high-efficiency cutting conditions.

To address these issues, the tool’s end edge shape is optimized by incorporating an arc design for edge protection. The concave straight edge is replaced with an arc edge where the tip is lower and the center is higher. The benefits include:

1.Increased cutting edge length, reducing the cutting load per unit, and distributing the cutting allowance and cutting force across the entire edge shape.

2.The bottom arc design, with a larger radius and smaller main cutting angle, reduces cutting forces and cutting-induced vibrations.

End Mill Cutter Shape Design and Analysis

Taking a φ10 four-flute end milling cutter as an example, the optimized arc-shaped milling cutter blade outperforms conventional end mills (corner radius and ball-end mills) in terms of effective cutting edge shape and length Le (black thick line) at the same cutting depth (0.5mm), as shown in Figure 1 and Table 1. The arc-shaped milling cutter has the longest effective cutting edge length, followed by the ball-end milling cutter, with the corner radius end milling cutter having the shortest. To comprehensively evaluate the performance of arc-shaped milling cutters compared to conventional end mills under high-efficiency machining conditions, both cutting simulations and cutting experiments were conducted for comparison.

During the cutting simulation, both the end milling cutter and the workpiece were simplified and precisely configured to ensure accurate calculations. A 10mm section of the end mill was selected as the simulation cutting portion, with detailed settings applied only to the cutting edge area. The total rotation angle of the tool during simulation was set to 190°, ensuring complete data for tool entry and exit points. The cutting parameters were set based on relatively large values commonly used in the mold industry: vc= 120m/min，fz=0. 4mm ／z，ap = 0. 5mm，ａe =10mm. The workpiece material selected was hardened steel (SKD11, with a hardness of HRC58), and the tool material chosen was carbide. The milling method used was climb milling. The cutting simulation model and the cutting conditions are shown in Figure 2.

Comparison of Cutting Forces

Cutting force is a crucial indicator of cutting performance. Excessive cutting force significantly impacts tool life. The cutting forces in the X, Y, and Z directions were directly extracted using the AdvantEdge post-processing program, showing how cutting forces fluctuate over time (see Figure 3).

From the figure, it is evident that the corner radius end milling cutter exhibits relatively stable cutting behavior. In contrast, the ball-end mill shows significant fluctuations in cutting force. This instability is attributed to the ball-end mill’s two-flute connection at the center, where long and short teeth alternate during milling. The variation in the number of active cutting edges leads to changes in the effective cutting edge length, resulting in substantial fluctuations in cutting force. The optimized arc-shaped end milling cutter initially encounters a larger cutting allowance when it begins to engage with the workpiece, resulting in higher cutting forces. As the cutting progresses deeper into the workpiece, the cutting allowance is uniformly removed radially along the tool, causing the cutting force to decrease and stabilize.

Comprehensive Comparison

The average values of the simulated cutting forces and cutting temperatures were calculated (see Table 1). It can be observed that the ball-end mill operates at a lower cutting temperature but experiences greater fluctuations in cutting force. The corner radius end milling cutter, with its shorter effective cutting edge, generates smaller cutting forces. The optimized arc-shaped end milling cutter produces higher cutting forces, predominantly in the axial direction.

Since the cutting edges involved in the process are concentrated on the end edges, they can be considered as simultaneously engaged in cutting. The resultant cutting force ftotalf_{total}ftotal? and the cutting force in the axial plane fxyf_{xy}fxy? per unit length of the cutting edge are shown in Table 2. The corner radius end mill exhibits the highest cutting force per unit length, indicating that under these conditions, the tool’s cutting edge experiences a higher cutting load, making it more prone to chipping and cutting vibrations. The optimized arc-shaped end milling cutter has the lowest cutting force per unit length, suggesting a more reasonable distribution of the cutting load.

Cutting Experiment Analysis

To comprehensively evaluate the cutting performance of the tools, the experiment was conducted to verify both cutting force and tool durability.

The experimental material, SKD11 with a hardness of HRC61, was the same as that used in the cutting simulation. The machining was performed on a MIKRON UCP1000 machining center using climb milling and dry cutting. The cutting process and parameters were consistent with those used in the cutting simulation.

Cutting Force Experiment

The cutting force sensor used was a Swiss Kistler 9265B three-component piezoelectric dynamometer, along with a charge amplifier and a corresponding data acquisition and processing system. After filtering, the cutting force values are shown in Table 3. The results indicate that the optimized arc-shaped end milling cutter generates a larger overall cutting force, but the cutting load per unit length is the smallest, consistent with the conclusions drawn from the cutting force simulation.

Cutting Performance Experiment for End Milling Cutter

Cutting hardened steel typically results in significant tool wear and short tool life, especially under high-efficiency cutting conditions, where differences in tool performance become more apparent. As shown in Figure 4, after machining a single groove (cutting length of 130mm), the optimized arc-shaped end mill exhibited normal wear, while the corner radius end mill and ball-end milling cutter both experienced edge chipping, leading to tool failure. The optimized arc-shaped end mill only showed edge chipping and failure after a cutting distance of 10 meters, which is more than ten times the tool life of conventional corner radius end mills and ball-end mills.

Analysis of Experimental Results

The experimental results indicate that cutting hardened steel leads to significant tool wear. Under high-efficiency cutting conditions, tools face extreme situations where only those that meet the machining demands can be used; otherwise, they are unsuitable. The optimized arc-shaped end milling cutter, with its modified blade shape, redistributes the cutting forces, reduces the cutting load per unit length on the effective cutting edge, and improves the tool’s cutting performance, thereby meeting the demands of high-efficiency cutting.

Conclusion

This study focused on modifying the tool blade shape to meet the high-efficiency cutting requirements for hardened steel. Through general cutting simulations and cutting experiments, the following conclusions were drawn:

1.Cutting simulations and cutting force experiments revealed that the optimized arc-shaped end mill generates greater cutting forces, particularly in the axial direction, compared to conventional corner radius and ball-end mills. However, it has a lower cutting force per unit length, with cutting temperatures comparable to those of the corner radius end mill.

2.Cutting experiments demonstrated that due to the modified blade shape, the cutting load per unit length of the cutting edge is reduced. As a result, the optimized arc-shaped end milling cutter outperforms conventional corner radius and ball-end mills under high-efficiency and heavy-load cutting conditions in mold manufacturing.

Experimental Materials and Methods

Tungsten carbide (WC) powders were used as the raw material. Four different particle sizes of WC powder were selected: 4.0 μm, 2.2 μm, 1.1 μm, and 0.5 μm, which were labeled as WC1, WC2, WC3, and WC4, respectively. Metal cobalt (Co) powder was used as the binder phase, and paraffin was used as the forming agent. Four different WC-Co mixtures with varying particle sizes were prepared.

The mixtures were processed into green bodies with specific shapes and densities using extrusion forming equipment. The pressed green bodies were then placed in a sintering furnace and subjected to high-temperature sintering at 1400°C for 30 minutes, followed by cooling, to form cemented carbide bars.

The magnetic properties of the cemented carbide bars were tested using a magnetic performance tester, measuring parameters such as coercive force (Hc) and saturation magnetization (Bs), and the results were analyzed.

Experimental Results and Analysis

Effect of Particle Size on Coercive Force

As shown in Table 1, the coercive force of cemented carbide bars increases with decreasing WC particle size, while the saturation magnetization also increases. This indicates that cemented carbide bars made with fine and ultrafine WC particles exhibit poorer magnetic properties. Among the samples, the ultrafine WC (WC4) shows the highest coercive force of 4450 A/m, followed by medium-sized WC (WC3) with a coercive force of 3300 A/m. Coarse WC (WC2) and very coarse WC (WC1) have lower coercive forces, at 2350 A/m and 1200 A/m, respectively. The increase in coercive force with decreasing WC particle size is primarily due to the increase in internal defects and dislocations within smaller particles. These defects and dislocations create resistance to domain wall movement, making the magnetization process more difficult and requiring a larger external magnetic field to achieve saturation, thereby increasing the coercive force.

Effect of Particle Size on Material Magnetic Performance Stability

For fine and ultrafine WC particles, the larger grain boundary area facilitates grain boundary diffusion and reactions, which reduces the material’s magnetic properties. As the WC particle size decreases, the magnetic saturation of cemented carbide bars gradually increases. Specifically: coarse WC (WC1) exhibits the lowest magnetic saturation at only 1.25 T; medium-sized WC (WC2) has a magnetic saturation of 1.15 T; fine WC (WC3) and ultrafine WC (WC4) show higher magnetic saturations at 1.05 T and 0.93 T, respectively. This is likely because fine and ultrafine WC particles have higher chemical reactivity, promoting the diffusion and bonding of the Co binder, thereby improving the stability of the material’s magnetic performance.

Magnetic saturation is an indicator of the remaining proportion of magnetizable material and is closely related to magnetic properties such as coercive force and remanence. The impact of WC particle size on magnetic saturation can be attributed to the degree of solubility of the binder phase in the cemented carbide bars. Coarse and medium-sized WC particles, having larger specific surface areas, have more contact with the Co binder, which enhances the solubility of Co in the cemented carbide bars. This effectively improves the material’s magnetic performance stability, resulting in higher coercive force and better magnetic stability. Conversely, fine and ultrafine WC particles, with smaller specific surface areas, reduce the effectiveness of the Co binder, potentially affecting the material’s hardness and magnetic properties. Thus, selecting the appropriate particle size during the preparation of cemented carbide bars is crucial for achieving the best overall performance based on specific application needs.

Impact of Gamma Phase on Material Performance

For cemented carbide materials, the proportion of the gamma phase directly affects the material’s hardness and magnetic properties. Variations in carbon and oxygen content also influence the gamma phase proportion and must be considered during material preparation. Generally, higher carbon content leads to an increase in the gamma phase proportion, thereby enhancing the material’s hardness and magnetic performance. Therefore, different WC particle sizes may have varying carbon and oxygen contents, which also affects the gamma phase proportion and the overall performance of the material.

Discussion on Sintering Atmosphere

In the sintering process of cemented carbides, the choice and control of the atmosphere have a decisive impact on the final microstructure and magnetic properties of the material. The atmosphere not only affects the chemical reactions during sintering but also directly relates to the microstructure and final performance of the cemented carbide. The types of sintering atmospheres are as follows:

Oxidizing Atmosphere:? air.

Reducing Atmosphere: Contains components such as H? or CO: hydrogen atmosphere for cemented carbide sintering.

Inert or Neutral Atmosphere: Argon, helium, vacuum.

Carburizing Atmosphere: Contains high components that cause carburization of the sintered body, such as CO, methane, and hydrocarbon gases.

Nitrogen-Based Atmosphere: High nitrogen content sintering atmosphere: 10% H? in N?.

We mainly selected vacuum, argon, and hydrogen atmospheres for discussion. The variations in coercive force and magnetic saturation of cemented carbides sintered in argon, vacuum, and hydrogen atmospheres differ depending on the atmosphere, as shown in Table 2.

From Table 2, it can be observed that under vacuum and argon atmospheres, the coercive force (Hc) of cemented carbide bar is higher compared to that in a hydrogen atmosphere. Conversely, the saturation magnetization (Bs) is lowest in a hydrogen atmosphere compared to vacuum and argon atmospheres.

Under vacuum and argon atmospheres, the effective control of oxygen partial pressure and the exclusion of volatile elements result in fewer pores and inclusions, clearer grain boundaries, and better grain growth, thereby enhancing the magnetic properties of the material. In contrast, in a hydrogen atmosphere, the reducing nature of hydrogen may reduce some elements in the cemented carbide, leading to the presence of uncertain phase components, poor grain growth, and subsequently affecting the material’s magnetic properties.

For coercive force (Hc), it is largely dependent on the material’s microstructure and magnetic anisotropy. Under vacuum and argon atmospheres, effective control of oxygen partial pressure and exclusion of volatile elements reduce magnetic anisotropy in the cemented carbide, which improves coercive force. However, in a hydrogen atmosphere, hydrogen’s reducing effect can lead to the reduction of some elements in the cemented carbide, resulting in grain defects and inclusions that directly affect magnetic anisotropy and reduce coercive force.

Regarding saturation magnetization (Bs), the relative magnetic saturation value in cemented carbide is influenced by factors affecting carbon content in the alloy. In vacuum or argon atmospheres, effective control of oxygen content reduces carbon loss. Although the pressed green body contains oxygen, which can be reduced by free carbon and carbon in WC (MeO + C = Me + CO), the oxygen content in these atmospheres is relatively low. In a hydrogen atmosphere, decarburization reactions (WC + 2H? → CH? + C) begin at around 100°C. Throughout the preparation process, the material is exposed to a decarburizing atmosphere, leading to a lower relative magnetic saturation value.

Conclusion

This experiment investigated the effects of different particle sizes and sintering atmospheres on the magnetic properties of cemented carbide bars. By comparing the magnetic properties of cemented carbide under different WC particle sizes (coarse, medium, fine, and ultrafine) and sintering atmospheres (vacuum, argon, and hydrogen), it was found that both particle size and atmosphere have a significant impact on the magnetic performance of the material.

From the perspective of particle size, as the WC particle size decreases, the coercive force of the cemented carbide bars increases, while magnetic saturation also increases. This indicates that particle size has a substantial effect on the magnetic properties of cemented carbide. Fine and ultrafine WC particles, due to their higher chemical reactivity and good sintering performance, can promote the diffusion and bonding of the Co binder, thus enhancing the stability of the material’s magnetic performance. However, smaller particle sizes may lead to increased porosity and inclusions, affecting the material’s hardness and magnetic performance. Therefore, the choice of particle size should be tailored to the specific application needs when preparing cemented carbide.

Regarding the atmosphere, cemented carbide bars sintered under vacuum and argon atmospheres exhibited higher coercive force and better magnetic stability. This is because these atmospheres effectively control the oxygen content and volatile elements, reducing porosity and inclusions, and promoting clearer grain boundaries and grain growth. In contrast, cemented carbide bars sintered in a hydrogen atmosphere showed significantly lower magnetic saturation. This is likely due to the decarburizing effect of hydrogen. Therefore, selecting the appropriate sintering atmosphere is crucial for obtaining cemented carbide bars with excellent magnetic properties. Further improvements in cemented carbide performance can be achieved by optimizing sintering process parameters and adding suppressants.

High-Speed Cutting Milling Equipment

1.High Stability of Machine Bed Components

The bed and support components of a high-speed cutting machine must exhibit excellent dynamic and static stiffness, thermal rigidity, and optimal damping characteristics. Most machines use high-quality, high-rigidity gray cast iron for these components, with some manufacturers incorporating high-damping polymer concrete into the base to enhance vibration resistance and thermal stability. This not only ensures stable machine accuracy but also prevents tool chatter during cutting. Measures such as closed bed designs, integral casting of the machine bed, symmetric bed structures, and dense ribbing are also crucial for enhancing machine stability.

2.Machine Spindle

The spindle performance of high-speed machines is crucial for achieving high-speed cutting. High-speed cutting spindles typically operate at speeds ranging from 10,000 to 100,000 RPM, with spindle power greater than 15 kW. Spindle axial gaps between the tool holder and spindle are controlled to be no more than 0.005 mm using compressed air or cooling systems. Spindles are required to have rapid acceleration and deceleration capabilities, meaning they must have high angular acceleration and deceleration rates.

High-speed spindles often use liquid static pressure bearings, air static pressure bearings, hot-pressed silicon nitride (Si3N4) ceramic bearings, or magnetic suspension bearings. Lubrication is commonly achieved with oil-air lubrication or spray lubrication, and spindle cooling typically involves internal water or air cooling.

3.Machine Drive System

To meet the demands of high-speed mould processing, the drive system of a high-speed milling machine should have the following characteristics:

High Feed Speed: Research indicates that increasing spindle speed and feed per tooth is beneficial for reducing tool wear, especially for small-diameter tools. Common feed speed ranges are 20-30 m/min, with large lead ball screws allowing speeds up to 60 m/min and linear motors achieving up to 120 m/min.

High Acceleration: High-speed milling of complex 3D surfaces requires a drive system with good acceleration characteristics, with drivers providing high-speed feed (fast feed rate of about 40 m/min and 3D profile processing speed of 10 m/min) and accelerations and decelerations ranging from 0.4 m/s2 to 10 m/s2.

Most machine manufacturers use closed-loop position servo control with small lead, large-size, high-quality ball screws or large lead multi-head screws. Advances in motor technology have led to the development and successful application of linear motors in CNC machines. Linear motor drives eliminate issues such as mass inertia, overshooting, lag, and vibrations, speeding up servo response, improving servo control accuracy, and enhancing machine processing precision.

4.CNC System

Advanced CNC systems are key to ensuring the quality and efficiency of high-speed processing of complex mould surfaces. Basic requirements for CNC systems in high-speed cutting include:

High-Speed Digital Control Loop: Includes 32-bit or 64-bit parallel processors and hard drives with over 1.5 GB; extremely short linear motor sampling times.

Speed and Acceleration Feedforward Control: Digital drive systems with jerk control.

Advanced Interpolation Methods: Such as NURBS-based spline interpolation for good surface quality, precise dimensions, and high geometric accuracy.

Look-Ahead Function: Requires a large capacity buffer register to pre-read and check multiple program segments (e.g., up to 500 segments for DMG machines, and 1000-2000 segments for Siemens systems) to adjust feed speeds and avoid over-cutting when surface shapes (curvatures) change.

Error Compensation Functions: Includes compensation for thermal errors due to linear motors and spindles, quadrant errors, measurement system errors, etc. Additionally, high data transmission speeds are required.

Data Interfaces: Traditional data interfaces like RS232 serial ports transmit at 19.2 kb, while many advanced milling centers now use Ethernet for data transmission at speeds up to 200 kb.

5.Cooling and Lubrication

High-speed milling uses coated carbide tools and operates without cutting fluids, resulting in higher cutting efficiency. This is because the high centrifugal forces of the rotating spindle make it difficult for cutting fluids to reach the cutting zone, and even if they do, the high temperatures may cause the fluids to evaporate, reducing cooling effectiveness. Additionally, cutting fluids can cause rapid temperature changes at the tool edge, leading to cracking. Thus, dry cutting with oil/air cooling is employed. This method quickly blows away the cutting heat with high-pressure air, and atomized lubrication oil forms a thin protective film on the tool edge and workpiece surface, effectively extending tool life and improving surface quality.

Tools for High-Speed Cutting

Tools are one of the most critical factors in high-speed cutting, directly impacting processing efficiency, manufacturing costs, and product precision. High-speed cutting tools must withstand high temperatures, pressures, friction, impact, and vibrations. They should have good mechanical properties and thermal stability, including impact resistance, wear resistance, and thermal fatigue resistance. The development of high-speed cutting tools has been rapid, with common materials including diamond (PCD), cubic boron nitride (CBN), ceramic tools, coated carbide, and titanium carbide (TiC) and titanium nitride (TiN) hardmetals.

For cutting cast iron and alloy steel, carbide is the most commonly used tool material due to its good wear resistance, although its hardness is lower than CBN and ceramics.

To improve hardness and surface finish, coating technologies such as titanium nitride (TiN) and aluminum titanium nitride (TiAlN) are employed. Coating technology has evolved from single-layer to multi-layer and multi-material coatings, becoming a key technology for enhancing high-speed cutting capabilities. Carbide inserts with titanium carbonitride coatings in the diameter range of 10-40 mm can process materials with Rockwell hardness below 42, while titanium aluminum nitride-coated tools can handle materials with Rockwell hardness of 42 or higher.

For high-speed cutting of steel, tools made from heat-resistant and fatigue-resistant P-class carbide, coated carbide, CBN, and CBN composite materials (WBN) are preferred. For cutting cast iron, fine-grain K-class carbide should be used for roughing, and composite silicon nitride ceramics or polycrystalline CBN (PCBN) tools for finishing.

For precision milling of non-ferrous metals or non-metallic materials, polycrystalline diamond (PCD) or CVD diamond-coated tools are recommended. When selecting cutting parameters, attention should be given to the effective diameter for round blades and ball end mills. High-speed milling tools should be designed with dynamic balancing, and the cutting edge angles should be adjusted compared to conventional tools.

High-Speed Milling Processes and Strategies

High-speed machining includes roughing, semi-finishing, finishing, and mirror finishing to remove excess material and achieve high-quality surface finishes and fine structures.

Roughing

The primary goal of mould roughing is to maximize material removal rate per unit time and prepare the geometric profile of the workpiece for semi-finishing. The process plan for high-speed roughing involves a combination of high cutting speeds, high feed rates, and small cutting depths. The most commonly used CAM software employs methods like spiral contouring and Z-axis contouring, which generate continuous, smooth tool paths in a single pass without retracting the tool, using arc entry and exit methods. Spiral contouring avoids frequent tool retraction and approach, minimizing the impact on surface quality and machine wear. Steep and flat areas are processed separately, with optimized tool paths generated using spiral methods with minimal retraction to achieve better surface quality. In high-speed milling, it is essential to use arc entry and exit methods and maintain a consistent tool path to minimize machine wear and achieve higher material removal rates.

Semi-Finishing

The semi-finishing process focuses on improving surface quality and dimensional accuracy, bridging the gap between roughing and finishing. The cutting speeds are higher than those used in traditional milling but lower than those in finishing. The primary goal is to achieve a better surface finish and precision by using a reduced depth of cut and controlling feed rates. Advanced CAM systems generate tool paths using techniques like trochoidal milling and adaptive clearing, which adaptively change cutting parameters based on the workpiece geometry and tool path. This method enhances tool life and surface quality while reducing cutting forces and thermal stresses.

Finishing

Finishing operations aim to achieve the final surface quality and dimensional accuracy. High-speed finishing involves higher cutting speeds and lower depths of cut, using techniques such as high-speed finishing cuts with constant engagement to ensure a smooth and uniform surface. Tool paths are optimized using advanced CAM software to achieve the desired surface finish and accuracy. Techniques like high-speed trochoidal milling and constant chip load milling are used to achieve excellent surface finishes and tight tolerances.

Mirror Finishing

Mirror finishing is the final step to achieve an exceptionally smooth and reflective surface. High-speed mirror finishing processes often involve special tools and techniques, including abrasive tools and polishing compounds. The key is to minimize surface irregularities and achieve a mirror-like finish with high precision. Techniques such as high-speed burnishing, polishing, and super-finishing are employed to achieve the desired surface quality.

Conclusion

High-speed milling technology has revolutionized the mould manufacturing industry by significantly enhancing machining efficiency, precision, and surface quality. The integration of advanced machining equipment, CNC systems, tooling technologies, and innovative milling strategies has enabled the production of complex mould cavities with high accuracy and reduced processing times. As technology continues to advance, high-speed milling will play an increasingly crucial role in meeting the evolving demands of the mould manufacturing industry.

Overview

In recent years, research on coarse grain carbide grades and materials has been advancing in two different directions: ultra-coarse and ultra-fine grains. Ultra-coarse grain cemented carbides have been widely applied in mining rock drilling tools, roll mills, and stamping molds.

Studies have revealed several primary forms of carbide failure during drilling: impact fatigue, abrasive wear, and thermal fatigue. For hard rock formations, such as granite (drilled with impact or rotary impact drills), abrasive wear is relatively lower, and carbide failure is primarily caused by impact and impact fatigue. The compressive strength and bending strength of the carbide are directly related to its impact fatigue resistance; additionally, this resistance is associated with the carbide’s purity, WC grain size, and Co phase’s average free path. Furthermore, the carbide’s impact fatigue resistance is directly related to the impact energy.

Reasons for Issues in Cemented Carbide Rock Drilling

For medium-hard rock formations, such as quartzite (drilled with impact drills), abrasive wear dominates. Abrasive wear generally consists of two aspects: micro-cracks at the contact points of abrasive particles and premature wear of the Co phase. The former primarily occurs on hard and brittle carbides, especially when abrasives have high fracture strength; the latter occurs on softer carbides with higher Co content, particularly when abrasives are very brittle. Figure 1 shows a scanning electron microscope (SEM) image of the wear surface of a GF20D grade drill tooth, produced by Xiamen Jinlu Special Carbide Co., Ltd., after drilling about 500 meters into quartzite. The YG6 grade carbide, composed of 94% WC with a grain size of 2-3 μm and 6% Co, has a hardness of HV30:1430. The image illustrates typical abrasive wear, characterized by premature Co phase wear and cracking and spalling of the WC phase.

For soft rock formations, such as sandstone, thermal fatigue is the primary cause of carbide failure, accompanied by abrasive wear. For ultra-soft rock formations, such as calcite and limestone, thermal fatigue is the main cause of carbide failure. The propagation of cracks and premature wear of the Co phase directly impact the drill tooth’s lifespan. Especially when drilling magnetite, thermal fatigue cracks, also known as creep cracks, dominate. Figure 2 shows a typical undulating cracking morphology of cemented carbide drill teeth formed while drilling magnetite. Figure 3 is an SEM image of a traditional polished cross-section of a carbide drill tooth that drilled about 5 meters into magnetite, composed of 94% WC with a grain size of 5 μm and 6% Co, with a hardness of 1230 HV. The image reveals that the thermal fatigue cracks on the carbide surface have extended into the carbide’s interior.

Figure 1: SEM photo of the wear surface of the quartzite at a depth of about 500 meters on the YG6 drill tips inserted in drill bits

Figure 2. Typical ups and downs of crack morphology formed when carbide drill teeth drill magnetite

Figure 3. The carbide drill teeth drill a conventional polished cross-section of about 5m into the magnetite. The grade consists of 94% WC with a grain size of 5um and 6% Co with a hardness of 1230HV(SEM).

Reasons for Developing New Rock Drilling Cemented Carbides

The fundamental reason for developing new rock drilling carbides lies in the continuous advancement of mining and drilling technology both domestically and internationally. As drilling equipment becomes more advanced and drilling efficiency improves, there is a growing use of fully hydraulic, high-power, and high-efficiency rock drilling rigs and rotary-percussion drills. This advancement has raised higher demands for the quality and lifespan of rock drilling cemented carbides. When drilling tools penetrate rock, the pressure rises from 0 to 10 tons within 1/10 of a second, and the temperature increases from 20°C to 1000°C. During impact and rotation, drilling carbides generate extremely high temperatures. Especially when drilling magnetite, rapid formation of thermal cracks, commonly referred to as “snake skin” or “tortoise shell” cracks, occurs.

To meet the requirements of modern rock drilling technology, the performance of rock drilling cemented carbides needs to be improved and optimized in several key areas: the thermal conductivity (the ability of the material to conduct heat) should be as high as possible; the thermal expansion coefficient (the linear expansion of the material when heated) should be as low as possible to ensure minimal growth rate of thermal cracks; high-temperature hardness should be further enhanced to guarantee good wear resistance at high temperatures; in addition, the transverse rupture strength (TRS) and fracture toughness (Kic, the material’s ability to resist sudden fractures caused by micro-cracks) should also be improved.

Table 1 lists the thermal performance data of pure WC, pure Co, three commonly used WC-Co carbide grades, and three types of rock. These three grades, with varying Co content and WC grain sizes, are suitable for different rock drilling teeth, hot-rolled rolls, and multi-purpose applications.

Directions for Developing New Rock Drilling Cemented Carbides

It is well known that Co has low thermal conductivity and a high thermal expansion coefficient. Therefore, the Co content should be minimized as much as possible. On the other hand, cemented carbides with high Co content exhibit better strength and fracture toughness. From a mechanical perspective, especially when carbide drill bits penetrate rock surfaces at high speeds, the drill bits endure high impact and loads, or mechanical vibrations under hard cutting conditions, necessitating improved strength and fracture toughness in the carbide. Additionally, compared to fine-grained carbides, coarse WC grain sizes contribute to greater strength and fracture toughness of the cemented carbide.

As a result, the preparation of rock drilling cemented carbides tends to use lower cobalt content and increase WC grain size to achieve good mechanical properties and the required high-temperature wear resistance. This approach results in ultra-coarse grain carbides. Traditionally, the production of ultra-coarse grain cemented carbides involves high-temperature reduction of coarse grain tungsten powder followed by high-temperature carburization to produce coarse grain WC powder. This powder is then mixed with Co powder and ball-milled to form a mixture, which is subsequently pressed and sintered to create the cemented carbide. However, coarse grain WC powder produced from tungsten powder via high-temperature carburization generally consists of polycrystalline particles, where each WC particle is composed of multiple WC single crystals.

Figure 4 shows a scanning electron microscope image of coarse grain carbide powder with a Feret diameter of 23.20 μm. The image reveals that each WC particle contains multiple WC single crystals. Although the original powder has a coarse grain size, after grinding, the polycrystalline particles easily break down into fine single crystal particles. Consequently, the ground WC powder has a Feret diameter of only 4.85 μm. Figure 5 shows the metallographic photo of a cobalt-containing carbide with 6% Co produced using conventional carbide production processes. The average grain size of this carbide is approximately 4.0 μm.

Figure 4: SEM image of coarse grain WC powder with a particle size of 23.20 μm.

Figure 5: Metallographic photo of WC-6% Co alloy produced from coarse grain WC powder with a particle size of 23.20 μm using conventional processing methods.

U.S. Patents 5505902 and 5529804 disclose methods for producing ultra-coarse grain cemented carbides. The methods outlined in these patents involve the dispersion and classification of coarse grain WC powder through jet milling and sieving to remove fine WC particles, selecting only the coarse-grained carbide, and then coating these WC particles with Co. Patent 5505902 utilizes the sol-gel method, where WC, methanol, and triethanolamine are mixed in a reactor. During heating, methanol evaporates, and Co precipitates onto the WC grains, forming a sol-gel.

Patent 5529804 employs the polyol method, where Co acetate, water, and WC are mixed and then spray-dried. The mixing process is optimized to prevent the breaking of coarse WC particles. The mixture produced using these patented methods is then subjected to conventional pressing and sintering processes to create cemented carbides with 6% Co and an average grain size of 13-14 μm, with porosity easily controlled between A02 and B02. This new carbide shows better WC matrix adjacency compared to carbides produced by traditional ball milling. Consequently, this new carbide has been successful in specific applications where conventional carbides fall short, such as in hard rock layers like granite and hard sandstone. In these cases, conventional column teeth fail due to Co dissolution at high temperatures, leading to spalling of elongated or hexagonal WC grains, and eventually, complete spalling of the drill bit within minutes, causing rapid crack propagation and subsequent fracture. In contrast, carbides produced with new technology can be used for extended periods in hard rock layers, displaying stable wear resistance without deep cracks. Due to the high adjacency of the WC matrix, the thermal conductivity of the 6% Co carbide with a WC average grain size of 14 μm can reach 134 W/m°C, which is 20% higher than that of coarse-grained carbides with the same Co content produced by traditional methods and comparable to the thermal conductivity of pure WC.

Application Examples of New Ultra-Coarse Grain Cemented Carbide Production Technologies

Two types of impact drilling cemented carbides were simultaneously produced using both traditional and new methods and tested in iron ore. Both samples had a WC average grain size of 8 μm, 6% Co, and 94% WC content.

Sample A: Produced using traditional ball milling, drying, pressing, and sintering processes. This carbide has a wide distribution of crystal sizes.

Sample B: The WC powder was subjected to jet dispersion and classification to remove coarser and finer WC particles, selecting 6.5-9 μm WC powder. The WC grains were pre-coated with 2% Co, and then 4% pure Co was added to achieve a 6% Co content. After wet mixing (without ball milling) to obtain the desired slurry, a thickening agent was added if necessary to prevent coarse grain WC sedimentation. The slurry was dried, shaped, and sintered, resulting in a narrower particle size distribution, with over 95% of the grains ranging from 6.5 to 9 μm. The adjacency of these carbides was measured: Sample A had an adjacency of 0.41, while Sample B had an adjacency of 0.61.

Testing was conducted in magnetite, which is prone to generating high heat and thermal fatigue. After drilling 100 μm, Sample A exhibited thermal cracking. Cross-sectional observation of the used carbide revealed small cracks extending into the carbide, damaging its microstructure and reducing its lifespan. With regrinding after every 100 μm of drilling, the carbide’s drilling lifespan was 530 meters. Sample B showed no or only minimal thermal cracking after drilling 100 meters. Cross-sectional observation showed no internal cracks, only some fractured surface grains. With regrinding after every 200 meters, the average drilling lifespan was 720 meters.

Structural Types

carbide?face milling cutters can be categorized into three types: integral welding type, mechanical clamping ?type, and indexable type.

The diagram 1 below illustrates an integral welding type face milling cutter. This type has a compact structure and is relatively easy to manufacture. However, if the teeth are damaged, the entire milling cutter must be discarded, so its usage has decreased.

As shown in the above diagram is the mechanical clamping welding type face milling cutter. This cutter welds carbide?inserts onto small cutter heads, which are then mechanically clamped into slots on the cutter body. When the inserts are worn out, they can be replaced with new ones, thereby extending the cutter body’s service life.

As shown in Figure 2, the commonly used indexable face milling cutter consists of components such as the cutter body (5), insert (1), tightening screws (3), cutter blade (6), wedge block (2), and eccentric pin (4). The insert (1) is clamped onto the cutter body using the wedge block (2) and tightening screws (3). Before tightening the screws, the eccentric pin (4) is rotated to adjust the axial runout of the insert within a specified range at the axial support point. Once the cutter blade (6) is mounted on the insert, it is clamped in place by the wedge block (2) and tightening screws (3). The eccentric pin (4) also prevents excessive axial forces on the insert during cutting, thereby preventing axial movement.

Compared to high-speed steel face milling cutters, carbide?face milling cutters offer higher milling speeds, better processing efficiency, and improved surface quality. They are capable of machining workpieces with hardened surfaces and layers, demonstrating significant advantages in enhancing product quality and processing efficiency.

Face Milling Cutter Main Structural Parameters

(1) Diameter and Number of Teeth

Diameter and number of teeth are the two main structural parameters of a face milling cutter. To accommodate different cutting requirements, face milling cutters of the same diameter are classified into coarse, medium, and fine types based on the number of teeth. Taking a 100 mm diameter cutter as an example, the number of teeth for coarse, medium, and fine types are 5 teeth, 6 teeth, and 8 teeth respectively.

(2) Geometric Angles

Indexable face milling cutters have key geometric angles including the lead angle κr, rake angle γp, and clearance angle γf. The lead angle κr is available in 45°, 60°, 75°, and 90° variants, with 75° being the most commonly used. When machining flat surfaces with shoulders or thin-walled workpieces, a 90° lead angle is typically chosen.

The rake angle γp and clearance angle γf can be combined into positive rake, negative rake, and positive-negative rake configurations. Positive rake angles are used for machining general materials; for instance, γp=7° and γf=0° are common for milling mild steel and cast iron, while γp=18° and γf=11° are used for milling aluminum alloys. Negative rake angles are employed for machining cast steel and hard materials, often set at γp=-7° and γf=-6°. Positive-negative rake angles offer good impact resistance and chip removal properties, suitable for milling general steel and cast iron, commonly used on machining centers with values like γp=12° and γf=-8°.

How to Select a Face Milling Cutter?

Selection of Face Milling Cutter Diameter

(1) When the machining area is not large, it is important to choose a tool or milling cutter with a diameter larger than the width of the plane. This allows for single-pass face milling. When the width of the face milling cutter is 1.3 to 1.6 times the width of the machining area, it effectively ensures proper chip formation and removal.

(2) For machining large surface areas, it is necessary to select a milling cutter with an appropriate diameter and perform multiple passes for face milling. Due to machine limitations, cutting depth, width, and the dimensions of the cutter and inserts, the diameter of the milling cutter may be constrained.

(3) When machining small plane areas or dispersed workpieces, a smaller diameter end mill should be selected for milling. To achieve optimal efficiency, the milling cutter should have contact with the workpiece equal to 2/3 of its diameter, which means the milling cutter diameter should be 1.5 times the width of the cut. Properly using this ratio of cutter diameter to cutting width ensures the milling cutter approaches the workpiece at an ideal angle. If the machine’s power cannot sustain cutting at this ratio, axial cutting thickness can be divided into two or more passes to maintain the ratio of cutter diameter to cutting width as much as possible.

Selection of Number of Teeth on the Milling Cutter

When selecting a milling cutter for machining, the number of teeth is an important consideration. For example, a coarse-toothed milling cutter with 6 teeth has a diameter of 100 mm, whereas a fine-toothed milling cutter with 8 teeth also has a diameter of 100 mm. The density of teeth affects both production efficiency and product quality. Dense teeth improve efficiency and quality but may hinder chip removal. Depending on the diameter of the teeth, they can be categorized as sparse teeth, fine teeth, and dense teeth.

Sparse teeth are used for rough machining of workpieces, with 1 to 1.5 inserts per 25.4 mm diameter, providing ample space for chips. Such tools are suitable for continuous chip formation in soft materials, using long blades and wide cuts. Dense teeth are advantageous for stable machining conditions, typically used for rough machining of cast iron, shallow and narrow cuts in high-temperature alloys, and when chip space is not required.

Dense teeth are applied in fine milling, with axial cutting depths ranging from 0.25 to 0.64 mm per tooth, minimizing cutting loads and power requirements, suitable for machining thin-walled materials.

Selection of Milling Inserts

The choice of milling inserts for flat milling is a critical factor to consider. In certain machining scenarios, pressed inserts are more suitable, while in others, ground inserts are preferred.

For rough machining

Pressed inserts are often preferred as they lower machining costs. Pressed inserts have lower dimensional accuracy and edge sharpness compared to ground inserts. However, they offer better edge strength, making them suitable for rough milling tasks. They can withstand higher impact and accommodate larger depths of cut and feed rates. Pressed inserts typically feature chip grooves on the front face, reducing cutting forces and friction with the workpiece and chips, thereby lowering power requirements. However, their surface finish is less compact than ground inserts, resulting in varying heights among insert tips on the milling cutter body. Due to their cost-effectiveness, pressed inserts find widespread use in production.

For fine milling

Ground inserts are preferable due to their superior dimensional accuracy. This high precision ensures precise positioning of the cutting edge during milling, leading to higher machining accuracy and lower surface roughness values. Moreover, the trend in ground milling inserts for fine machining includes forming large positive rake cutting edges with chip grooves, allowing the inserts to handle small feed rates and depths of cut effectively. In contrast, carbide?inserts without sharp rake angles may experience friction with the workpiece during fine machining with small feed rates and depths of cut, reducing tool life.