Experimental Method

TiN, TiAlN, TiN-MoS?, and CrTiAlN composite coatings were deposited on YT14 cemented carbide cutters using a closed-field unbalanced magnetron sputtering ion plating equipment from Teer Company. The nano hardness and elastic modulus of the coatings were measured using a Nano Test 600 nano hardness tester. To reduce experimental errors, the hardness and elastic modulus values were the average of five measurements. The hardness of the coatings was also verified using a Vickers microhardness tester, and the morphology and phase structure of the cutter coatings were observed and analyzed using a Quanta 200 scanning electron microscope (SEM) and an Advance 8 X-ray diffractometer (XRD). The cutting tests of the coated cutters were conducted on a CNC machining center, with PCNiMoVA steel as the cutting material. The flank wear was observed and measured using a 30x tool microscope. The cutting time when the wear strip width on the flank face exceeded 0.6mm was used as the basis for evaluating the tool’s life, and the cutting life of the tools was compared.

Experimental Results and Analysis

Coating Hardness and Elastic Modulus Testing

Figure 1 shows the loading and unloading curves obtained during the nano hardness measurement process of the CrTiAlN composite coating. The loading and unloading curves not only provide the hardness of the CrTiAlN film but also its elasticity. Define R = (ha – h) / h as the elastic recovery coefficient, where ha is the indentation depth at maximum load, and h is the residual depth of the indentation after unloading. According to the definition of R, the greater the value of R, the greater the elasticity of the film. Therefore, based on the nano indentation curve in Figure 1, the hardness of the CrTiAlN film is 33 GPa, and the elastic modulus is 675 GPa. Figure 2 is a comparative analysis chart of the nano hardness of TiN, TiAlN, TiN-MoS?, and CrTiAlN coatings. It can be seen from the figure that the nano hardness measurement values of the four coatings are 18 GPa, 30 GPa, 15 GPa, and 33 GPa, respectively. The nano hardness ranking is as follows: CrTiAlN > TiAlN > TiN > TiN-MoS?.

Figure 3 shows the measured elastic modulus values of each coating. It can be seen from the figure that the elastic modulus of the four coatings are 214 GPa, 346 GPa, 164 GPa, and 675 GPa, respectively. The ranking of the elastic modulus is as follows: CrTiAlN > TiAlN > TiN > TiN-MoS?. This indicates that the elastic modulus of the coatings is directly proportional to their hardness. However, the CrTiAlN coating shows the greatest relative increase in elastic modulus, with a value significantly higher than other coatings, reaching 675 GPa, which suggests that the deposited CrTiAlN coating has both high hardness and high elasticity.

Meanwhile, the Vickers microhardness tester was used to perform microhardness verification tests on each cutter coating, with a indentation load of 15g and a duration of 10 seconds. The measurement results are shown in Figure 4. By comparing the coating nano hardness in Figure 2 with the coating microhardness values in Figure 4, it can be found that the microhardness variation trend of each coating is the same as the nano hardness variation trend, with the CrTiAlN coating having the relatively highest Vickers microhardness at HV1560.

Analysis of Coating Surface Morphology

The surface morphologies of the TiN, TiAlN, TiN-MoS?, and CrTiAlN coated cutters are shown in Figure 5. It can be observed from the figure that there is a significant difference in the surface morphology of the four coatings, indicating that the addition of composite elements has caused a great change in the crystalline state of the TiN compounds. Among them, the TiN coating surface phase tissue is uniform, with relatively fine particles, while the TiAlN coating surface morphology is relatively rough, with coarser particle tissue. The TiN-MoS? coating surface is distributed with a large amount of flaky mixed structure, mainly due to the uniform distribution of MoS? phase in the TiN coating, tending towards a composite structure in the mixed state, which serves a self-lubricating function. The CrTiAlN coating surface grains are relatively fine, the coating is dense and uniform, and the surface is distributed with a large number of hard points.

Drilling Test

The four types of coated cemented carbide cutters were used to machine PCNiMoVA steel, and the wear condition of the cutters was inspected to compare the durability of different coated cutters.

The cutting test conditions for the coated cutters were as follows: the cutting method was external cylindrical cutting, the cutting speed was 160 m/min, the feed rate was 0.15 mm/r, the cutting depth was 0.5 mm, and dry cutting was performed. The cutting time when the wear strip width on the flank face exceeded 0.6 mm was used as the basis for evaluating the tool’s life, and the cutting life of the tools was compared.

A comparison of the cutting life of the coated cutters is shown in Figure 6.

From the figure, it can be seen that under dry cutting conditions, the cutting life of the uncoated cutters was the shortest, and the service life of the coated cutters was significantly better than that of the uncoated cutters. Among them, the CrTiAlN coated cutter had the longest cutting life. The ranking of the cutting life of the four coated cutters was as follows: CrTiAlN > TiN-MoS? > TiAlN > TiN. This indicates that Cr and Al elements form hard phases in the TiN coating, and the addition of Al elements is beneficial for the formation of Al oxides, which avoids further oxidation during the cutting process, improves the oxidation resistance of the cutters, and is conducive to increasing the cutting life of the tools. Meanwhile, the MoS? lubricating phase helps to reduce the friction coefficient of the cutters, improve the anti-wear capability of the tools, and also extends the service life of the cutters.

In summary, due to the comprehensive utilization of the advantages of various coating components in the multi-component composite coatings, they achieve better comprehensive performance, ensuring excellent wear resistance and toughness, reducing the formation of built-up edge, and possessing mechanical shock and thermal shock resistance, which can greatly improve the tool life.

XRD Test Analysis

XRD analysis method was used to characterize the phase structure of the CrTiAlN tool coating with the best cutting performance, and the results are shown in Figure 7. The XRD spectrum analysis indicates that the crystal phases of the coating are mainly composed of Cr, CN, CEN, and TiN at room temperature, and the amorphous phase in the coating was not detected. At the same time, high magnification scanning analysis of the tool coating revealed a large number of hard phase particles distributed on the coating surface. Combined with X-ray diffraction analysis, it is known that these hard phases are mainly CN, CrN, TiN, and AlN phases. These hard phases are beneficial for improving the cutting life of the coated cutters.

Conclusion

The author has prepared TiN, TiAlN, TiN-MoS?, and CrTiAlN composite coatings using a closed-field unbalanced magnetron sputter ion plating PVD coating process. Comparative tests on the mechanical properties and cutting performance of the coatings show that:

The nanoindentation analysis obtained the nano hardness ranking of the four types of cutter coatings as follows: CrTiAlN > TiAlN > TiN > TiN-MoS?. The elastic modulus of the coatings is directly proportional to their hardness.

Under dry cutting conditions, when drilling PCNiMoVA steel, the cutting life of the coated cutters is ranked as: CrTiAlN > TiN-MoS? > TiAlN > TiN. This indicates that the cutting performance of the multi-component composite coatings is significantly better than that of the simple TiN coating, suggesting a direction for the future development of coated cutters.

Main Preparation Techniques for Micro-Nano Cemented Carbide Powders:

1.Low-Temperature Reduction Decomposition Method

This method is an improved version of the conventional reduction carburization method to prepare micro-nano cemented carbide powders. It uses a low-temperature hydrogen process to reduce tungsten acid, tungsten oxide, or tungstic acid to micro-nano-scale tungsten powder, followed by carburization of the tungsten powder to micro-nano-scale WC powder at low temperatures.

2.Mechanical Alloying (MA)

This is a traditional method using mechanical force for the chemical synthesis of micro-nano cemented carbide powders. It involves placing a certain proportion of elemental powder mixtures in a ball mill jar and subjecting them to high-energy ball milling under an inert atmosphere for a long time. The powder particles undergo repeated grinding, breaking, extrusion, cold welding, and low-temperature solid-state chemical processes under the action of mechanical force, resulting in alloy powders with uniform composition and structure.

3.Spray Drying Method

Also known as thermochemical synthesis, this is currently the main method for industrial mass production of WC-Co composite powders. The process was developed by L.E. McCandlish and B.E. Kear of Rutgers University and has been patented. The Nanodyne Company in the United States uses this process to produce nanoscale WC-Co powder with a particle size of 20～40nm. The process involves mixing ammonium metatungstate [(NH?)?(H?WO??O?o)·4H?O] aqueous solution with cobalt chloride (CoCl·nH?O) to form an original solution, which is then atomized and dried to form a uniformly composed, fine mixture of tungsten and cobalt salts, followed by reduction and carburization in a fluidized bed to obtain nanophase WC-Co powder.

4.Gas Phase Reaction Method

This method uses the principle of gas-phase chemical reaction deposition to produce powders. It involves evaporating and vaporizing metals or alloys in equipment and reacting with active gases at certain temperatures to produce metal compounds, which are then condensed to obtain micro-nano-scale compound powders.

Experimental Method

High-purity cobalt powder and 0.8μm tungsten carbide powder were used as raw materials, mixed in a WC-10%Co ratio, and subjected to high-energy ball milling with a ball-to-material ratio of 9:1 using a QM-IF type planetary ball mill. The ball milling time was set at 24 hours per interval, with milling times of 24, 48, 72, and 96 hours. The particle size of the milled powder was measured using the Fsss method and X-ray diffraction. The powder samples with different ball milling times were then pressed and sintered under the same conditions to prepare cemented carbide specimens. Subsequently, strength and hardness tests were conducted, along with metallographic analysis.

Results and Analysis

The Influence of Ball Milling Time on Powder Particle Size and Grain Size

Table 1 shows the measured Fsss particle sizes. It can be seen that the powder particle size decreases with the extension of ball milling time but becomes coarser after decreasing to a certain extent. However, the grain size continuously decreases. Powder particles and grains are different concepts. Particles consist of multiple grains encapsulated by cobalt powder. As ball milling progresses, while powder particles break, cold welding can also occur between the cobalt on the particle surfaces. Therefore, when ball milling reaches a certain degree, particle agglomeration exceeds breaking, leading to an increase in particle size, which eventually maintains a certain equilibrium state. The measured grain size is usually WC grains, which are brittle phases and easily broken during ball milling. Due to the encapsulation by cobalt powder, it is difficult for WC grains to cold weld, so the grain size continuously decreases.

Microstructure After Sintering

Figure 3 shows the metallographic structure obtained after sintering, clearly indicating the influence of ball milling time on the grain size, shape, and distribution. As the ball milling time extends, the grains in the sintered body are significantly refined, tend to be uniform in size, and the WC grains are more dispersed. This is due to two factors: first, the WC grains themselves are refined and homogenized through ball milling, indicating that ball milling not only breaks the WC grains but also homogenizes them; second, it is formed during the sintering process. The sintering process only coarsens the particles and causes non-uniformity due to abnormal grain growth, which typically becomes more severe with longer ball milling time, as ball milling can cause lattice distortion, promoting abnormal grain growth. However, the results of this experiment do not show this; instead, the grain size tends to be more uniform with the extension of ball milling time. Clearly, the homogenization of WC grains is due to the effect of ball milling. This demonstrates that high-energy ball milling can refine and homogenize WC grains.

The Effect of Ball Milling Time on the Mechanical Properties of Ultrafine Cemented Carbide

Tables 1 and 2 show the measured bending strength and hardness, respectively. It can be seen from the tables that both strength and hardness increase with the extension of ball milling time. As previously analyzed, the longer the ball milling time, the finer the WC grains in the sintered samples and the more uniform their distribution. This indicates that bending strength and hardness increase simultaneously with the refinement of grains. In the grain size range above the micron level, the strength and hardness of cemented carbide typically have an inverse relationship; that is, as bending strength increases, hardness decreases, and vice versa. However, in this case, both have improved simultaneously, which clearly shows that after high-energy ball milling, the obtained cemented carbide has reached the micro-nano scale grain size range. The hardness improvement of the cemented carbide in this study is very significant; generally, WC-10%Co has a hardness of about HRA91, but here it has reached as high as 92.8. This indicates that grain refinement has a very strong strengthening effect on cemented carbide.

Conclusion

This paper has conducted a preliminary study on the relationship between ball milling time and grain size in WC-Co ultrafine cemented carbide, as well as the relationship between grain size and strength, hardness, and the following conclusions are drawn:

High-energy ball milling has a very strong breaking effect on WC grains, and the WC grain size refines with the increase of ball milling time. However, there is a critical point for the powder particle size during ball milling. Upon reaching this critical point, the grain size of the cemented carbide is the smallest, after which, with the increase of ball milling time, the grains may become coarser instead.

Micro-nano grain size cemented carbide can be obtained through high-energy ball milling, and the WC grains are more uniform and dispersed.

The refinement of the grain size of cemented carbide to the micro-nano scale can simultaneously increase the bending strength and hardness.

This paper aims to perform high-speed machining of TC4 titanium alloy using a vibration-damping end mill (end mill with unequal tooth pitch angles). The MATLAB software is used to perform a Fast Fourier Transform (FFT) on the milling force, with a focus on analyzing the impact of chatter on the machining of titanium alloys. The objective is to optimize the cutting speed while ensuring the quality of the machined surface and low cutting force, thereby improving cutting efficiency.

Dynamic Milling Force Model

To study the cutting stability, a reasonable dynamic model needs to be established. Compared to the workpiece with higher stiffness, the end mill can be considered as an elastic body. Since the end mill has a very high stiffness in the axial (z-direction) direction, the milling machining system can be simplified into a “spring-damping” system with two mutually perpendicular degrees of freedom in the x-direction (feed) and y-direction, as shown in Figure 1. Φj(t) represents the rotation angle of the end mill; hi(t) represents the dynamic chip thickness; fz is the feed per tooth.

Establishing the equations of motion for 2 degrees of freedom?movement:

In the formula: m, s, c, k represent the mass, displacement, damping coefficient, and elastic coefficient of the cutting system, respectively; Fx(t) and Fy(t) are the dynamic milling forces in the x and y directions, respectively. The milling forces studied in this paper are all Fy, and the motion equation is simplified to a single degree of freedom motion equation:

The milling force Fy(t) is related to the dynamic chip thickness hi(t) during the milling process and is a function of the end mill’s rotation angle φj(t), making it a periodic function. The variation frequency of the cutting force is the tool tooth passing frequency (TPF). Due to manufacturing and clamping errors causing the tool system to be asymmetric, its variation frequency is the spindle rotation frequency (SF).

The spindle rotation frequency (SF) is defined as:

In the formula: ω is the angular velocity of the spindle rotation, in radians per minute; v is the linear velocity of the spindle, in meters per minute; D is the diameter of the end mill, in millimeters; the unit of spindle rotation frequency is Hz.

The tool tooth passing frequency (TPF) is defined as:

In the formula: N represents the number of teeth on the end mill.

Milling Experiment

Experimental Setup, Workpiece Material, and Cutting Tool

The milling experiment setup is shown in Figure 2. The machine used is a DAEWOO ACE-V500 machining center with a rated power of 15 kW and a maximum torque of 286.2 N?m. The milling force is measured by a Kistler 9257B three-coordinate force measuring instrument with a sampling frequency of 7,000 Hz. The cutting tool used is a 4-fluted variable pitch solid carbide end mill, and the distribution of the tooth pitch angles at the bottom is shown in Figure 3. The diameter of the end mill is 20 mm, and the overhang length is 74.3 mm. The workpiece material is TC4 titanium alloy.

Section 2.2: Experimental Design

The experiment was conducted using dry cutting and conventional milling. The cutting speed v, spindle rotation frequency (SF), and tool tooth passing frequency (TPF) are listed in Table 1. The feed per tooth fz, axial cutting depth ap, and radial cutting depth ae were kept constant at 0.08 mm/tooth, 20 mm, and 0.5 mm, respectively. The range of cutting speed varied from 80 to 360 m/min. Figure 4 shows the milling sequence, where the sequence numbers correspond to those in Table 1.

Experimental Results

Time-Domain Analysis of Cutting Forces

Figure 5 shows the relationship between the maximum amplitude of Fy and the cutting speed (v). From the figure, it can be observed that when the cutting speed is in the range of 80~160 m/min, the maximum milling force remains essentially unchanged. When the cutting speed reaches 200 m/min, the milling force suddenly increases, and when the cutting speed reaches 240 m/min, the milling force reaches its first peak. Subsequently, as the cutting speed increases,The cutting force significantly decreases until the cutting speed increases to 320 m/min, at which point the milling force reaches its minimum. When the cutting speed is 360 m/min, the milling force reaches its maximum value.

Analysis of Milling Force Spectrum

The frequency domain provides more information about the cutting process than the time domain. As the direct responder of the tool and workpiece, the milling force can describe the vibration condition of the cutting system. From the collected y-direction milling force Fy signal, a 4-second segment of data in the middle was selected and processed using MATLAB software for Fast Fourier Transform (FFT) to analyze the milling force spectrum and determine the machining condition.

Since the variation frequency of the milling force is TPF, the peaks of the milling force generally appear at TPF and its integer multiples (n·TPF) in the frequency domain. However, due to tool manufacturing and mounting errors, the variation frequency of the milling force is often SF, meaning the spectrum peaks of the milling force appear at SF and its integer multiples (n·SF).

Figures 6(a) to (h) show the milling force spectrum analysis at different cutting speeds. From the figures, it can be seen that when the cutting speed is 80 and 120 m/min, the peaks of the milling force spectrum appear at the spindle rotation frequency (SF) and the tool tooth passing frequency (TPF). When the cutting speed is 160, 280, and 320 m/min, the peaks of the milling force appear not only at SF and TPF but also at high frequencies. The peak at the spindle rotation frequency (SF) is larger at higher cutting speeds than at lower cutting speeds, which is due to the mass imbalance caused by tool eccentricity. As the cutting speed increases, the centrifugal force increases, leading to an increase in the y-direction milling force peak. When the cutting speed is 240 and 360 m/min, the larger peaks at high frequencies do not appear at SF and its integer multiple frequencies, indicating that chatter occurs in the cutting system at these times, with chatter frequencies of 734 Hz (v=240 m/min), 730 Hz, 1,111 Hz, 1,268 Hz, 1,363 Hz, and 1,459 Hz (v=360 m/min). Combined with Figure 5, it can be determined that when chatter occurs, the radial milling force increases significantly.

Machined Surface Quality

The Wyko NT9300 white light interferometer was used to measure the machined surface topography. This instrument uses the principle of optical interference and can achieve nanometer precision in topography measurement.

Figure 7 shows the machined surface topography measured by the white light interferometer. From the figure, it can be seen that when v=160 m/min, the surface topography is the best. When v=240 and 360 m/min, there is a large area of material removal on the machined surface. This is partly due to chatter, which increases friction, extrusion, and tearing between the tool and the workpiece, causing the tool to produce unstable vibrations on the workpiece surface, leading to overcutting in some areas. On the other hand, these violent movements cause the cutting temperature to rise, resulting in adhesion between the flank of the tool and the workpiece. As the tool moves, most of the adhered workpiece material is carried away by the tool.

By comparing Figure 5 with Figure 8, it is found that the roughness curve follows a similar trend to the cutting force curve. When chatter occurs (at cutting speeds of 240 and 360 m/min), both the cutting force and roughness reach their maximum values, proving that chatter not only increases the cutting force but also increases the surface roughness.

From Figure 8, it can also be seen that for stable cutting (cutting speeds of 80 to 160 m/min), the higher the cutting speed, the lower the surface roughness. Although stable cutting occurs at cutting speeds of 200 and 280 m/min, Figures 6(d) and (f) show that the high-frequency peaks of the milling force are significant, which affects the surface quality. When the cutting speed is 320 m/min, the cutting force is small, and the spectrum peaks are also small, but the surface roughness is poor. This is due to the tool eccentricity (larger peak at SF) having a significant impact on surface roughness at high cutting speeds.

Conclusion

Under dry cutting conditions, when high-speed milling TC4 titanium alloy, chatter occurs at cutting speeds of 240 m/min and 360 m/min, with chatter frequencies around 730, 1,111, 1,363, and 1,459 Hz.

For stable cutting, the cutting speed has little effect on the magnitude of the milling force. However, when chatter occurs, the cutting force significantly increases, and the surface quality of the machining deteriorates.

At a cutting speed of 160 m/min, the surface roughness is low, and the cutting force is small, so 160 m/min can be recommended as the optimal cutting speed.

Research on the grinding process of WC-based cemented carbide mainly focuses on grinding speed, grinding time, and ball-to-material ratio, with less research on the shape of the grinding media. Therefore, this paper selects grinding media of different shapes, prepares WC-10%Co cemented carbide, and studies the influence of the shape of the grinding media on the micro-morphology of the powder, the morphology, and performance of the alloy, thereby providing a reference for the development of a reasonable grinding process.

Experimental Methods and Procedures

The technical parameters of the WC powder used are shown in Table 1. Wet grinding was carried out using a cemented carbide wet mill jar. WC powder, Co powder, and Cr?C? were proportioned according to the experimental design plan in Table 2, with a ball-to-material mass ratio of 4:1. Different WC-Co grinding media (spherical 06.5 and rod-shaped 07×14, as shown in Figure 1) were selected for wet grinding and mixing in a drum mill at 90 r/min. The addition of wet grinding medium alcohol was 280 mL-kg1, the binder was PEG4000 (2.0 wt.%), and the grinding time was 25h and 40h. After grinding, the slurry was placed in a vacuum drying oven to dry. After screening, the mixture was pressed into green bodies of 6 mm×10 mm×15 mm specification at a pressure of 150 MPa. The green bodies were dewaxed and sintered in a sintering furnace. Dewaxing was carried out using a slight positive pressure, with a temperature range of 180-500°C. The sintering temperature was 1450°C, with a holding time of 2h, and finally, the cemented carbide samples were obtained.

The micro-morphology of the samples was observed using a German Zeiss EVO 18 scanning electron microscope, and the average grain size of the alloy was measured using the intercept method. The specific surface area of the powder was determined using an American Conta Monosorb MS specific surface area meter. The density of the samples was obtained by the Archimedes method, and the relative density was calculated. The coercive force (Hc) of the samples was tested using a coercive force analyzer (Zhongda-ZDHC40). The transverse rupture strength (TRS) of the samples was tested according to the GBT 3851-1983 B standard using a Sansi UTM5105 electronic universal testing machine. The fracture toughness of the samples was tested according to the ASTM B771 standard using a Sansi UTM5105 electronic universal testing machine. The hardness of the samples was tested using a Rockwell hardness tester (Wilson-RS74).

Analysis and Discussion of Results

Figure 2 shows the SEM images of powders prepared with different shapes of grinding media and different grinding times. As shown in the figure, with the increase of grinding time, the average particle size of the powder gradually decreased whether using grinding balls or rods. This is because as the grinding time increases, the breaking and extruding effects of the grinding media and the mill jar on the powder also deepen continuously. The more energy produced during grinding, the more intense the impact and shear the powder receives, leading to the generation of a large number of dislocations. The particles continuously break along the particle interfaces and grain boundaries, resulting in the continuous refinement and homogenization of the powder.

The state of cemented carbide powder during time ball milling less than 40 hours

When the grinding time reached 40h, the specific surface area of the powder obtained by rod grinding was 2.01 m2·g1, which was higher than the 1.85 m2·g1 obtained by ball grinding. The finer the powder, the larger the specific surface area, and the higher the powder activity, with greater surface energy, making it easier to agglomerate together and adsorb oxygen. This is beneficial for pore shrinkage and the disappearance of vacancy clusters during sintering, achieving densification.

Powder status after ball milling time more than 40h

After the grinding time reached 40h, the powder produced using grinding balls had some coarse particles and generated more broken powder, which adhered to the surface of larger particles or agglomerated together. Additionally, the powder produced using grinding rods appeared rounder in appearance, while the powder produced using grinding balls had an irregular shape. This is due to the different contact methods of the two shapes of grinding media, as shown in Figure 3 for a schematic of the contact between grinding balls and rods. The contact method between grinding balls is point contact, which easily produces a larger force at the points of contact, leading to a higher likelihood of breakage. Moreover, during the movement of the grinding balls, the contact with the powder is non-selective, resulting in low precision of the breakage. This leads to the powder ground with balls being irregularly broken, producing a large amount of broken powder. In contrast, the contact method between grinding rods is a combination of line contact and point contact. During the grinding process, the force applied at the points of contact is more dispersed, avoiding the generation of large forces and thus preventing over-grinding. The grinding rods have a selective breaking action that breaks coarse particles while protecting fine particles. In the grinding process, the coarse particles are necessarily the first to be ground, making the probability of coarse particle powder being ground higher than that of fine particle powder. This results in the powder ground with rods being more uniform.

Figure 4 shows the SEM images of alloys prepared with different shapes of grinding media and different grinding times. It can be observed from the figure that WC grains are randomly distributed in the Co phase, with shapes typical of irregular rectangles and triangles . As the grinding time reaches 40h, the size of the WC grains is reduced to varying degrees, and the cemented carbide grains obtained with grinding rods are finer and more uniform. In contrast, the alloy grains obtained with grinding balls exhibit a significant issue of coarse grain inclusion, with obviously large grains present. It can also be found in Figure 4(b) that there are a large number of pores near the large grains, and the presence of large grains and numerous pores will inevitably affect the mechanical properties of the alloy. Figure 5 shows the particle size distribution of alloys prepared with different shapes of grinding media and different grinding times. After the grinding time extends from 25h to 40h, the average grain size (D) of the spherical medium decreases from 1.530 μm to 0.618 μm, and the average grain size of the rod-shaped medium decreases from 1.847 μm to 0.538 μm. The distribution curve in the graph becomes narrower, and the standard deviation significantly decreases, indicating that the grains become more uniformly distributed as the grinding time increases. Combined with Figure 4, it can be found that after 40h of grinding, the alloy grains obtained with grinding rods are finer and more uniform, while the cemented carbide grains obtained with grinding balls have coarse grain inclusion, with obviously large grains observable. This is because the grinding intensity of the grinding balls is higher than that of the grinding rods, which easily produces broken powder and irregular large particles, leading to abnormal grain growth during sintering.

Table 3 shows the properties of cemented carbides prepared with different shapes of grinding media and different grinding times. After 25h of grinding, the relative density of the alloy ground with grinding balls is 97.4%, which is higher than the 97.1% of the alloy ground with grinding rods; when the grinding time is increased to 40h, the relative density of the alloy ground with grinding rods is 99.6%, higher than the 99.0% of the alloy ground with grinding balls. The main factors affecting the relative density of cemented carbide include pores, specific surface area, carbon content, and composition. As the grinding time is extended, the relative density improves, which could be due to the alloy grains becoming finer and more uniform with the extended grinding time, resulting in reduced pores and increased density. After 40h of grinding, the alloy ground with grinding balls has large grains, and as observed in Figure 4(b), there are a large number of pores near the large grains, which weakens the density. The alloy ground with grinding rods has finer and more uniform grains, and the powder has a larger specific surface area and higher activity, which is beneficial for pore shrinkage and the disappearance of vacancy clusters during sintering, promoting densification and thus achieving a higher relative density.

It can be seen from Table 3 that as the grinding time extends, the hardness and bending strength of the alloys ground with both types of grinding media increase. At 40h of grinding time, the hardness and bending strength of the alloy ground with grinding rods are higher. First, due to the dispersion and mixing of agglomerated mixtures during grinding, as the grinding time increases, the WC grains become finer, and the finer WC grains will reduce the contact between each other, prompting an increase in the average free path of the Co phase, with a more uniform distribution, increasing the effective deformation range, and thereby increasing the hardness and bending strength of the cemented carbide . The alloy ground with grinding rods for a longer time has finer and more uniform grains, with a better fine-grain strengthening effect, so at 40h of grinding time, the hardness and bending strength of the alloy ground with grinding rods are higher. Secondly, hardness and bending strength are also closely related to density and pores; the presence of pores will weaken the alloy’s ability to resist damage, according to the empirical formula:

In the formula: σ represents the strength of the cemented carbide corresponding to the porosity P; σ0 represents the strength of the alloy when the porosity is zero; b is a constant; P is the porosity. Under the same conditions, the higher the porosity of the material, the smaller the effective area that bears the load, resulting in a lower corresponding material strength. Therefore, as shown in Figures 4 and 3, the alloy ground with grinding balls has large grains, and there are a large number of pores near the large grains, which reduces its density. The alloy ground with grinding rods has finer and more uniform grains, fewer pores, and greater density. Hence, as the grinding time increases, the hardness and bending strength of the alloy are enhanced, with the alloy ground with grinding rods being higher.

Finally, according to the Hall-Patch formula, the relationship between the alloy strength and grain size is as follows:

In the formula, σ represents the strength of the cemented carbide ; d represents the grain size. Combining Figures 2 and 4, when the grinding time reaches 40h, the grains ground with rod-shaped media are finer and more uniform, and the density is also higher. Therefore, the more refined the alloy grains become, the stronger the fine-grain strengthening effect will be, leading to a higher bending strength and hardness of the alloy. In contrast, the alloy prepared at a grinding time of 25h, especially when using grinding rods, results in coarser grains, lower density, and more pores. This leads to a lower average free path of the Co phase, uneven distribution of the binder phase, and a shorter effective deformation range. Consequently, the bending strength and hardness of the alloy are both low.

For the fracture toughness of the samples, it can be observed from Table 3 that at a grinding time of 40h, the fracture toughness of the alloy ground with rod-shaped media is 9.5 MPam1/2, which is lower than that of the cemented carbide ground with spherical media at 10.3 MPam1/2, and this is inversely proportional to the hardness relationship between the two. Combining with Figure 4, it can be found that the alloy ground with grinding balls may have coarse grains, which hinder crack propagation, making transgranular fracture more likely to occur at large grains. Moreover, the larger the grain size, the stronger the ability to accommodate moving dislocations, and the greater the resistance to crack propagation. The cemented carbide ground with grinding rods has finer and more uniform grains, mainly exhibiting intergranular fracture, making crack propagation easier and leading to a decrease in fracture toughness. Therefore, the fracture toughness of the alloy ground with grinding balls is higher than that of the alloy ground with grinding rods.

It can also be known from Table 3 that as the grinding time increases, the coercive force of the alloy ground with grinding balls increases from 103 kA m-1 to 123 kA m-1, and the coercive force of the alloy ground with grinding rods increases from 97 kA m-1 to 129 kA m-1. According to literature reports, the coercive force of the alloy has the following relationship with the WC grain size and Co content:

In the formula, Hc represents the coercive force of the alloy; dwc represents the grain size of WC; Wc represents the mass fraction of Co in the alloy. In this experiment, the Co content is the same for all samples. Therefore, it can be understood that the thinner the grain size of the alloy, the smaller the thickness of the magnetic bonding phase will be, and the more evenly it will be dispersed, leading to an increase in coercive force. When the grinding time reaches 40h, whether using grinding balls or rods, the coercive force of the cemented carbide is significantly improved due to the refinement of the grains with increased grinding time. Since the alloy ground with grinding rods has finer and more uniform grains, whereas the alloy ground with grinding balls has larger grains and poorer uniformity, the coercive force of the alloy ground with grinding rods is higher.

Conclusions

WC-10%Co cemented carbide with the same composition was prepared using grinding media of different shapes. The micro-morphology of the powder and the morphology and properties of the alloy were studied and analyzed, and the following conclusions were drawn:

Compared to grinding for 25h, the powder ground for 40h with both shapes of grinding media is more refined. However, the grinding intensity of the grinding balls is higher than that of the grinding rods, leading to the presence of coarse grains and broken powder in the ground material, which affects the properties of the alloy.

The grain size of the alloy ground for 40h with both shapes of grinding media is finer and more uniform than that ground for 25h. Compared to the alloy ground with grinding balls, which has abnormally large grains deteriorating the grain distribution and properties of the alloy, the alloy ground with grinding rods has finer and more uniform grains, enhancing the properties of the alloy. The alloy ground with grinding rods for 40h can achieve better properties: relative density of 99.6%, coercive force of 129 kA·m-1, hardness of HRA 91.5, fracture toughness of 9.5 MPa·m1/2, and bending strength of 3565 MPa.

How to Solve Burr Issues

Chemical Deburring

Chemical deburring utilizes chemical energy for processing. Chemical ions adhere to the surface of the parts, forming a film with high electrical resistance and low conductivity, which protects the workpiece from corrosion. Since metalworking burrs are higher than the surface, they can be removed through chemical action. This method of deburring is widely used in fields such as pneumatic, hydraulic, and construction machinery.

High-Temperature Deburring

First, place the parts that need deburring into a tightly sealed chamber, then introduce the chamber into a hydrogen-oxygen mixed gas environment with a certain pressure. Ignite the mixed gas to cause an explosion, releasing heat to burn off the burrs without damaging the parts.

Roll Grinding Deburring

Place the parts and abrasive materials into a closed drum. As the drum rotates, a dynamic torque sensor, along with the parts and abrasive, generates a grinding action to remove metalworking burrs. Abrasives can be made of quartz sand, wood chips, alumina, ceramics, and metal rings, among others.

Manual Deburring

This method is more traditional and also the most time-consuming and labor-intensive. It mainly involves manually grinding with tools such as steel files, sandpaper, and grinding heads. The most commonly used tool in production now is the trimming knife.

Process Deburring – Edge Rounding

Edge rounding can refer to any action that removes the sharpness from the edges of metal components. However, it is typically associated with creating a radius on the edges of parts.

Edge rounding is not simply about removing sharpness or deburring, but about breaking the edges of metal components to improve their surface coating coverage and protect them from corrosion.

Removing metalworking burrs in Milling Parts

In milling parts, deburring is more complex and costly, as multiple metalworking burrs are formed at different positions with varying sizes during milling. It is particularly important to select the correct process parameters to minimize metalworking burr size.

Main Factors Affecting Burr Formation in End Milling

① Milling parameters, milling temperature, and cutting environment, among others, have a certain impact on burr formation. Some key factors, such as feed rate and milling depth, are reflected through the plane cutting angle theory and the Edge Engagement Sequence (EOS) theory.

② The better the plasticity of the workpiece material, the more likely it is to form Type I burrs. In the processing of brittle materials with end milling, if the feed rate or the plane cutting angle is large, it is conducive to the formation of Type III burrs (deficits).

③ When the angle between the end face of the workpiece and the machined plane is greater than a right angle, the increased support stiffness of the end face can suppress machining burr formation.

④ The use of cutting fluid is beneficial for extending tool life, reducing tool wear, lubricating the milling process, and thereby reducing the size of burrs.

⑤ Tool wear has a significant impact on machining burr formation. When the tool wears to a certain extent and the tip radius increases, not only does the burr size in the tool retracting direction increase, but type burrs may also form in the tool cutting-in direction.

⑥ Other factors, such as tool material, also have a certain impact on burr formation. Under the same cutting conditions, diamond tools are more effective in suppressing metalworking burr formation than other types of tools.

How to Effectively Handle Burrs Generated During Tool Retraction

To suppress burrs generated during tool retraction, eliminating the space where burrs are produced is an effective method. For example, measures such as chamfering can be taken to reduce the space before retracting the tool.

Using appropriate cutting conditions to suppress machining burrs should aim to minimize the amount of cutting residue, and it is necessary to select the most suitable tool and cutting conditions. Use tools with a large rake angle and sharp cutting edges. Increase cutting speed to improve cutting characteristics. Especially during finish cutting, it is essential to use the minimum cutting depth and feed rate.

The size of the space between the tool and the workpiece determines the size of the burrs. Let’s take a look at the relationship diagram below.

Summary

In fact, during the machining process, burrs are inevitable, so it is best to address the metalworking burr issue through process improvements, avoiding excessive manual intervention. Using chamfering end mills can reduce the space where burrs are produced, effectively remove burrs, and is also a very suitable method for clearing burrs.

Common Failure Phenomena of Tungsten Carbide Roll Rings

Sand Holes Inside the Roll Ring

The production process of tungsten carbide cemented carbide roll rings generally includes hot pressing, cold pressing, hot isostatic pressing, and cold isostatic pressing to achieve the densification of the internal structure of the roll ring and ensure its wear resistance and sufficient strength. However, due to various factors during the production process, sand holes similar to those produced by casting processes can occur within the roll ring’s structure, leading to failure in use. When slotting or grinding the roll rings, we often encounter this phenomenon. The sand holes vary in size; smaller ones can be eliminated by increasing the grinding amount during the roll ring processing, and their elimination does not affect the use of the carbide roll ring. Larger sand holes cannot be removed by general grinding and sometimes require abandoning a particular roll groove during processing. If the sand hole is located exactly in the middle of the roll ring’s width and is large, the roll ring can only be scrapped.

Cracks at the Bottom of the Roll Ring Groove

Under normal production, each rolling groove can guarantee a certain amount of rolling. However, some carbide roll rings wear too quickly during the rolling process and fail long before reaching the rated rolling amount. During inspections, it can be found that some roll rings have regular nail-like microcracks at the bottom of the rolling groove (see Figure 1); while others may have a craze pattern. These two types of cracks are dangerous for carbide roll rings. If the roll is not replaced in time, the cracks will expand, leading to the scrapping of the carbide roll ring.

Roll Ring Bursting

During the production process, there are two phenomena of roll ring bursting: circumferential fracture and radial fracture. In circumferential fractures, cracks occur at the bottom of the roll ring’s axial hole grooves and spread in a ring shape along the rolling groove; in radial fractures, the roll ring cracks radiate radially. In production practice, circumferential fractures of roll rings are relatively rare, with most occurrences being radial fractures (see Figure 2).

Roll Ring Shattering

Shattering of roll rings is another major phenomenon of roll ring failure and is a more serious accident that occurs during the rolling process. The hazard of roll ring shattering is extremely high because during the rolling process, the rolling mill operates at high speed, and the explosive fragments of the roll ring can damage other roll rings, leading to the expansion of the accident. When a carbide roll ring shatters, fragments often strike and injure the carbide roll rings of adjacent frames, damaging the taper sleeve and aluminum cap.

Analysis of Roll Ring Failure Causes

Cracking Due to Repeated Thermal Stress

During the rolling process, the hot rolled piece comes into contact with the surface of the rolling groove, causing the temperature of the roll surface to rise. This part of the metal will expand, while the metal temperature in the deeper layers of the roll rises less, resulting in compressive stress on the surface metal of the roll. Conversely, when the roll surface is quenched by cooling water, the surface metal contracts, while the deeper metal does not contract as much as the surface metal, creating tensile stress in the surface layer. This repeated alternation of thermal stress easily generates thermal fatigue cracks, causing the roll ring groove bottom to appear nail-like microcracks and crazing.

Optimizing Cooling Water Pressure

Insufficient cooling water pressure is also one of the causes of carbide roll ring fatigue cracks. To reduce the cracks caused by fatigue, it is necessary to use cooling water to carry away the heat obtained from the rolling piece, thereby reducing the temperature rise of the roll ring and the thermal expansion of the surface metal. When the rolling piece comes into contact with the roll ring surface, the surface metal of the carbide roll ring can reach 500~600°C. When cooling water is sprayed onto the hot roll ring surface, it forms a layer of steam film that covers the underlying roll surface, severely affecting the cooling effect. Studies have shown that when the pressure of the cooling water is less than 0.5MPa, it cannot break through the steam film, and even with sufficient water volume, the desired cooling effect cannot be achieved.

Maintaining Water Quality

Water quality can have a significant impact on the life of the finishing mill roll rings. Tungsten carbide roll rings have special requirements for the acidity of the water quality. Generally, acidic water quality can cause corrosion of the hard alloy rolls. When the pH value of the water is below 7, the cobalt alloy begins to corrode, exacerbating the propagation of thermal cracks and greatly reducing the life of the roll rings. Therefore, in addition to maintaining cleanliness, the acid-base balance of the water must also be maintained.

Quality, Operation, and Production Factors

The circumferential fracture of the roll ring is mainly caused by the propagation of fatigue cracks, while the reasons for the radial radiating fracture of the roll ring are more complex. Summarizing practical experience, the following are several causes of radial fracture of carbide roll rings:

Material Deviations and Sintering Process Defects

There are issues with the quality of the roll ring itself. The material formulas of roll rings produced by different manufacturers are distinct, resulting in different grades of roll rings. Even carbide roll rings of the same grade produced by the same manufacturer may have deviations in material proportions, leading to unstable quality of the roll rings. Another reason is that during the manufacturing process of the roll rings, improper control of the sintering process parameters can lead to defects in the quality of the carbide roll rings. We have experienced continuous radial fractures of roll rings provided by a domestic manufacturer, and the manufacturer admitted that there were issues with the sintering process after analyzing the causes of the accident together.

Overlooked Process Requirements and Excessive Pressure

Improper operation during roll mounting. There are clear process requirements for the mounting and dismounting of finishing mill roll rings, but in the production process, operators often overlook these requirements in order to work faster, resulting in insufficient cleaning of the roll rings and the taper sleeves. Moreover, operators may not control the pressure properly when using the roll mounting cart, and excessive pressure can easily cause roll ring fractures during the rolling process. Additionally, if the temperature of the roll box shaft head is not sufficient when the operator mounts the roll, it can also easily cause roll ring fractures during the rolling process.

Taper sleeve quality issues

Because the inner surface of the taper sleeve contacts the roll shaft of the roll box, and the outer surface contacts the inner surface of the roll ring, the process requires high precision in terms of concentricity and ovality. If the concentricity and ovality are out of tolerance, uneven contact between the taper sleeve and the roll ring can cause local stress during production, leading to carbide roll ring fractures.

Steel stacking

It is inevitable that steel stacking accidents occur during production. If such an accident occurs in the finishing mill, it can cause significant damage and harm to the roll rings, as the rolling pieces are rolled at high speed with temperatures above 1000°C. When steel stacks, the metal accumulates around the roll ring groove, and due to high temperatures, the metal adheres to the roll ring surface, forming a “turtle shell” phenomenon. At this point, the local heating of the roll ring generates thermal stress, easily forming thermal fatigue cracks, which can lead to fracture failure when the roll ring continues to be used.

Conclusion

The reasons for roll ring shattering are also multifaceted. Generally, it is an expansion of the roll ring fracture accident, as the roll ring fractures while the mill is still operating at high speed, and the tremendous centrifugal force causes the fractured roll ring to break apart. Another reason is the impact of the rolling piece on the roll ring

Mechanism of Tool Breakage

Abrasive Wear by HardParticles

This is mainly caused by mechanical wear from impurities in the workpiece material, hard particles such as carbides, nitrides, and oxides contained in the material matrix, and fragments of built-up edge. These hard particles scratch grooves on the tool surface. Abrasive wear by hard particles occurs in tools at all cutting speeds, but it is the main cause of wear in low-speed steel tools. At low cutting temperatures, other forms of wear are not significant. It is generally believed that the amount of wear caused by abrasive wear by hard particles is proportional to the relative sliding distance or cutting distance between the tool and the workpiece.

Adhesive Wear

Adhesion refers to the bonding phenomenon that occurs when the tool and workpiece material come into contact at atomic distances. It is a so-called cold welding phenomenon that occurs due to plastic deformation on the actual contact surface of the friction surfaces under sufficient pressure and temperature. It is the result of the adhesive force between the fresh surface atoms formed by the plastic deformation of the two friction surfaces. The adhesive points on the two friction surfaces are sheared or stretched and carried away by the opposite surface due to relative motion, which causes adhesive wear.

Adhesive wear can occur on the contact surfaces between two materials, whether on the soft material side or the hard material side. Generally, the breakdown of adhesive points occurs more frequently on the side with lower hardness, i.e., the workpiece material. However, the tool material often has defects such as uneven structure, internal stresses, microcracks, pores, and local soft spots, so the carbide tool surface also frequently breaks and is carried away by the workpiece material, forming adhesive wear. High-speed steel, ceramic, cubic boron nitride, and diamond tools can all experience wear due to adhesion. The size of the carbide grains in cemented carbide has a significant impact on the speed of adhesive wear. The temperature at which the tool material and workpiece material adhere to each other greatly affects the severity of adhesive wear. Other factors such as the hardness ratio of the carbide tool to the workpiece material, the shape and structure of the tool surface, as well as cutting conditions and the stiffness of the process system, all affect adhesive wear.

Diffusion Wear

Due to the high temperatures during cutting, and the fact that the carbide tool surface is constantly in contact with the freshly cut surface, which has high chemical reactivity, the chemical elements of the two friction surfaces may diffuse into each other. This results in a change in the chemical composition of both, weakening the properties of the tool material and exacerbating the wear process. Diffusion wear increases with the rise in cutting temperature. For a given tool material, as the temperature increases, the rate of diffusion initially increases slowly and then accelerates. Different elements have different diffusion rates, so the severity of diffusion wear is greatly related to the chemical composition of the carbide tool material. Additionally, the rate of diffusion is also related to the flow velocity of the chip layer on the tool surface, which corresponds to the velocity of the chip flowing over the rake face. Slower flow velocity results in slower diffusion.

Figure 1 shows the diffusion wear of WC-Co cemented carbide, indicating that tungsten carbide (WC) and cobalt (Co) have dissolved into the surface layer of the steel, and this surface layer has also melted at the interface. In the figure, the upper layer is steel, the lower layer is cemented carbide, and the middle white layer is the melted layer, which is in a locally melted area, with WC grains surrounded within it. Because the temperature is higher at the crescent-shaped depression on the rake face, the diffusion rate is high and wear occurs quickly. At the same time, since adhesion occurs when the temperature rises to a certain degree, diffusion wear and adhesive wear often occur simultaneously, easily forming a crescent-shaped depression.

Chemical Wear

Chemical wear occurs at certain temperatures where the tool material reacts chemically with surrounding media (such as oxygen in the air, extreme pressure additives like sulfur and chlorine in cutting fluids, etc.), forming a layer of compounds with lower hardness on the tool surface, which is then carried away by the chips, accelerating tool wear; or because the tool material is corroded by a certain medium, causing tool wear.

The main types of normal tool wear include hard particle wear, adhesive wear, diffusion wear, and chemical wear, and there are interactions among them. For different tool materials, under different cutting conditions, and when machining different workpiece materials, the primary cause of wear may be one or two of these types.

Methods to Improve Tool Performance

Because it is difficult to balance the wear resistance and toughness of cemented carbide tool materials, users can only select suitable tool materials from various grades of cemented carbide based on specific machining objects and conditions, which brings inconvenience to the selection and management of cemented carbide tools. To further improve the comprehensive cutting performance of cemented carbide tool materials, current research focuses include the following aspects:

Grain Refinement

By refining the grain size of the hard phase, increasing the intergranular surface area of the hard phase, and enhancing the intergranular bonding strength, the strength and wear resistance of cemented carbide tool materials can be improved. When the WC grain size is reduced to sub-micron levels, the hardness, toughness, strength, and wear resistance of the material can all be enhanced, and the temperature required for full densification can also be reduced. The grain size of conventional cemented carbide is 3~5μm, fine-grained cemented carbide has a grain size of 1~1.5μm (micron level), and ultra-fine-grained cemented carbide can have a grain size below 0.5μm (sub-micron, nano-level). Common grain refinement processes include physical vapor deposition, chemical vapor deposition, plasma deposition, and mechanical alloying. Since these grain refinement processes are not yet mature, nano-grains easily grow into coarse grains during the sintering of cemented carbide, and the general growth of grains will lead to a decrease in material strength. Individual coarse WC grains are often a significant factor causing material fracture. On the other hand, fine-grained cemented carbide is relatively expensive, which also restricts its promotion and application.

Surface, Overall Heat Treatment, and Cycle Heat Treatment for Carbide tools

Surface treatments such as nitriding and boriding on the surfaces of cemented carbide with good toughness can effectively improve their surface wear resistance. For cemented carbide with good wear resistance but poor toughness, overall heat treatment can change the bonding composition and structure in the material, reduce the adjacency of the WC hard phase, thereby improving the strength and toughness of the cemented carbide. Using cycle heat treatment processes to alleviate or eliminate stress at the grain boundaries can comprehensively improve the overall performance of cemented carbide materials.

Addition of Rare Metals

Adding rare metal carbides such as TaC and NbC to cemented carbide materials can form complex solid solution structures with the original hard phases WC and TiC, further strengthening the hard phase structure. At the same time, it can inhibit the growth of hard phase grains and enhance the uniformity of the structure, which is greatly beneficial to improving the comprehensive performance of cemented carbide. In the standard P, K, M grades of cemented carbide, there are grades that have added Ta(Nb)C (especially more in the M grade).

2 The forming process and main characteristics of boundary notch in cemented carbide cutting tools

The boundary notch of cemented carbide cutting tools is a wear area,which is?relatively large,resulting from friction between main cutting edge and the surface of?the workpiece as the following Fig.1.Fig.1(a)shows a traditional wearing type of the?flank.The rake face A,and flank face Aa?are also shown.Fig.1 (b)shows the main?dimension of boundary notch of the lathe tool,in which VN represented the height of?boundary notch and C refers to the width.It is apparent that the greater the?dimensions of VN and C are,the greater it destroys the performance of tools and influences the machining quality.

By experiment,the forming process of the boundary?notch can be divided into the?following three steps:firstly,several micro cracks are produced at main cutting edge.Secondly,the mesh fractures are found in the boundary areas and they will spread.

Finally,the piece material will be denuded and the boundary notch is formed.In the?subsequent cutting process,the dimension of the boundary becomes bigger and?bigger.

Fig.2 shows the forming process of boundary notch of the cemented carbide?cutting?tools.

Main factors to influence boundary notch are mechanical performance of the?piece material,the cutter material,and geometry parameter of the cutter.The?following experiments were carried out in order to expound the forming mechanism?and evolution rules of the boundary notch.

3 Experiment conditions and testing measures for assessing boundary notch in cemented carbide cutting tools

The lathe C6130 and reversible cutting tool are used in the experiment.Five?cutter materials are employed.Main mechanical parameters of cutter material are?shown in Table 1.

The machining piece is the friction-welded line of the single hydraulic pillar.The?width of the welded line is 15mm and the machining allowance is 5.5mm.Besides,the above pillar is welded with 270SiMn and 45#steel.The relatively mechanical?performances of the welded line are shown in Table 2.

Based on manufacturing experience and relative information in China and other?countries about similar machining process,the chosen machining and tool geometry?parameters are shown in Table 3.

The boundary notch dimensions of the cemented carbide cutting tools (boundary?notch height VN and width C are directly attained by tool microscope.In order to?ensure reliability of the results,repeated experiments are carried out.The recurrent?performance is good.

Influences of Cutting angles

Influences of Cutting Edge Angle

The results of the variety boundary notch are shown as in Fig.4 when the cutting?edge angle is changed.From Fig.4 we can find that,with the lessening of the cutting?edge angle Kr,the dimensions of the boundary notch decrease.The reason is that?with the lessening of the cutting edge angle Kr,the length of the cutting edge that?acts on cutting becomes larger and the average loads on the cutting edge be?come?lighter.

Influences of Cutter Corner Radius r

The results of the variety boundary notch with the cutter corner changing are?shown as Fig.5.The boundary notch dimension decreases with the cutter corner radius?are becoming lesser.The reason is that with the increasing of the cutter corner radius,the impact-resistance performance.

Therefore,under the same cutting conditions,boundary notch dimensions (VN,C)decrease?when the cutter corner radius becomes lesser.

Influences of Negative Chamfer bal in cutting tool

The experiment results of the variety boundary notch are shown as in Fig.6 when?the width of the negative chamfer is changed.The dimension of the boundary notch?will decrease when the width of the negative chamfer ba decreases.Therefore,in?order to resist or decrease the cutter boundary notch,the lesser negative chamfer bal should be chosen.

Deburring Machining Process

The burrs have some influences on cutter boundary notch in metal machining?process.A deburring cutter is chosen to decrease the adverse influence on cutter.A?different result between deburring machining process and common machining process?is shown as in Fig.7.It can be seen that about 75%of the boundary notch is?decreased.So,burr is a main factor to produce and increase the boundary notch of the?cutter.

Conclusion

From above experimental research and theoretical analysis,the following?conclusions are attained:

1)Boundary notch of the cutting tool can be expressed by boundary notch height?VN and boundary notch width C.The forming processes of boundary notch can be?divided into three steps:micro-tipping appears firstly;Then,mesh fractures expand;Finally,boundary notch results.

(2)Main factors that influence boundary notch of cemented carbide cutter are?piece material,cutter material and cutter geometry parameters.

(3)Deburring machining process and adjusting cutting tool geometry parameters(to reduce edge angle K,and width of negative chamfer ba,to increase cutter corner?radius re)can be chosen to decrease effectively boundary notch,which ensures the?quality of workpiece and cutting?performances of cutting tool.

Development of Grades

Due to the combination of coating technology and the carbide matrix with high toughness and high hardness, the market segmentation of standard K and P type carbides is changing, and the application range of K type carbides is still expanding. The development process of extruded grades is as follows.

K10/20, with a cobalt content of about 6%, contains approximately 0.2% VC and 1% TaC, and the grain size of WC is between 1.0 and 1.5 μm. Since large TaC grains (>2 μm) are prone to premature shedding during cutting, and also reduce the toughness of the carbide, which is the biggest taboo for K-type carbides, the TaC in this grade carbide?has been replaced by Cr?C?. Subsequently, the K20F grade appeared, with a cobalt content of about 8%, and the WC grain size reached 0.7 μm, mainly used for PCB micro-drills and twist drills. When extended to cutters for paper processing, the cobalt content increased to 8.5%, and the WC grain size was 1.0 to 1.5 μm. Jinlu Company’s GK10, GK20, and GK20UF correspond to these grades. The typical microstructures of GK10, GK20, and GK20UF grade carbides are shown in Figures 1, 2, 3, and 4.

ISO grade K30-K40, with a cobalt content of 10%, also adds Cr?C? and VC. The WC grain size is between 0.6 and 0.8 μm. This grade of carbide?has high strength (TRS reaching 4000 MPa), high hardness (HRA 92.2), and high wear resistance. It is important to note that the total content and the ratio of Cr?C? and VC have a significant impact on the hardness and strength of the carbide, and for different processing objects, the content and proportion of Cr?C? and VC need to be selected. To this day, this grade is still one of the main grades for rods. Jinlu Company’s GK30F grade corresponds to this. The typical microstructure of the GK30F grade carbide?is shown in Figures 5 and 6.

Japanese manufacturers were the first to develop hard metal drills and end mills with a cobalt content of 12%, while also adding Cr?C? and VC (totaling 1.2%), with a WC grain size of 0.4 μm. The wear resistance and toughness of this grade of carbide?are significantly improved, thereby significantly extending the tool life. It has outstanding advantages in the processing fields of hardened steel, stainless steel, titanium carbides, and glass fiber-reinforced plastics. The processing performance advantages of this grade are becoming more apparent and it is replacing the GK30F grade in many application areas. Jinlu Company’s GK40UF grade corresponds to this. The typical microstructure of the GK40UF grade carbide?is shown in Figures 7 and 8.

The ultra-fine grade currently under development, which contains a cobalt content of 12% and a WC grain size of 0.2 μm to 0.3 μm, and also contains VC and Cr?C? inhibitors, is widely favored by the industry. The microstructure of the test sample for Jinlu Company’s GK30SF grade is shown in Figure 9.

Tables 1 and 2 list the composition and performance indicators of the aforementioned grades.

Development of Production Technology

Production Method of Extrusion Materials

The production process of extrusion materials is as follows: wet grinding (both rolling ball milling and stirring ball milling can be used) → slurry drying → blending with a forming agent → screening and granulating. It must be noted that a considerable number of extruded product grades contain one or two types of trace grain growth inhibitors, therefore the selection of WC powder and cobalt powder grades and the method of adding grain growth inhibitors will fundamentally affect the performance of the carbide.

Extrusion Methods

Currently, the extrusion machines used in the production of hard metal extrusions include plunger extruders and screw extruders.

Plunger extruders’ output per working cycle is limited by the volume of the plunger cavity. Different tonnage plunger extruders have varying single loading quantities, ranging from a few kilograms to several dozen kilograms. Their advantage is flexible production scheduling, convenient switching between different grades of materials, and minimal material loss.

Screw extruders have almost unlimited production capacity, with continuous feeding and extrusion, which is highly suitable for the mass production of established grades. The downside of screw extruders is that equipment cleaning and installation are time-consuming when switching grades, and there is a significant waste of materials.

Both plunger extruders and screw extruders, advanced extrusion technology manufacturers have made progress in the automatic placement and automatic cutting of the compact. The extrusion molding of porous synchronous extrusion and double helix cooled bar stock is a reflection of the overall advancement of extrusion technology (see Figures 10 and 11).

Method of Removing Forming Agents

The composition of the extrusion forming agents and the corresponding process for removing them are topics that manufacturers have continuously researched. The motivation for this research comes from reducing costs and improving the dimensional limits of extruded molding. The traditional process involves removing the forming agent in a hydrogen atmosphere, followed by vacuum low-pressure sintering. Technologically advanced manufacturers have made breakthroughs in forming agent research, with the extrusion blank first being dried in an oven (without the need for a protective atmosphere) before undergoing vacuum low-pressure sintering. The drying process and equipment are shown in Figures 12 and 13.

Sintering Technology

Extruded profiles all use vacuum low-pressure sintering processes to ensure the uniformity of the product’s material and the stability of production.

Inspection and Performance Evaluation

In addition to the conventional analysis indicators of the carbide, the quality inspection of extruded profiles also includes ultrasonic testing for the sintered blanks of rods with a diameter of more than 12mm. Attention should be paid to the dispersion of the product’s bending strength and the fluctuation of the coercive force. For micro-drill rod materials used in PCBs, special attention must be given to controlling the carbide’s hardness, bending strength, and magnetic saturation value; all three indicators must be controlled simultaneously.

Conclusion

Just like with die-pressed cemented carbide products, there is a significant gap between China’s extruded cemented carbide profiles and the world’s advanced level. The main aspects of this gap are as follows:

• Material quality disparities
• Disparities in binder and pre-sintering technology
• Inability to produce, or produce stably in bulk, cutting-edge products
• Automation disparities
• Stability of industrial production (within the same batch or between different batches)

Therefore, we should increase our research and development efforts in the aforementioned areas to quickly narrow the gap with the world’s advanced level. This will provide reliable domestically produced cutting tools for China’s emerging electronic industry and machinery manufacturing industry.

Three-Axis Machining Center

The three-axis machining center is the most widely used, including the X, Y, and Z axes, also known as three-axis simultaneous machining centers. The three-axis machining center can perform simple plane machining, but it can only machine one side at a time. It can effectively machine materials such as metal, aluminum, wood, etc.

Functions and Advantages of Three-Axis Machining Centers

The most effective machining surface of a vertical machining center (three-axis) is the top surface of the workpiece. A horizontal machining center, with the aid of a rotating table, can only complete the machining of the four sides of the workpiece. Currently, high-end machining centers are developing towards five-axis control, allowing the completion of five-sided machining in one setup. With the configuration of a high-end five-axis simultaneous CNC system, high-precision machining of complex spatial surfaces can also be achieved.

Four-Axis Machining Center

The four-axis machining center adds an additional rotation axis to the three-axis, usually the A-axis. The rotation of the A-axis allows the workpiece to rotate around the vertical axis on the horizontal plane, enabling multi-face machining. The four-axis machining center is suitable for situations where machining is required on different faces of the workpiece, such as inclined surfaces, oblique holes, etc.

Features of Four-Axis Simultaneous Machining

(1) Machining that cannot be achieved by three-axis simultaneous machining machines or requires overly complex fixturing.

(2) Improvement of the precision, quality, and efficiency of free-form surfaces.

(3) The difference between four-axis and three-axis is the addition of one rotation axis. The establishment of the four-axis coordinates and their code representation:

Determination of the Z-axis: The direction of the machine spindle axis or the vertical direction of the workpiece fixture is the Z-axis.

Determination of the X-axis: The horizontal plane parallel to the workpiece mounting surface or the direction perpendicular to the workpiece’s rotation axis within the horizontal plane is the X-axis, with the direction away from the spindle axis being the positive direction.

Five-Axis Machining Center

The five-axis machining center adds another rotation axis to the four-axis, usually the C-axis. The rotation of the C-axis allows the workpiece to rotate around an axis perpendicular to the table, enabling more complex multi-angle and surface machining. The five-axis machining center is suitable for complex shapes, multi-angle machining, including spatial surface machining, special-shaped machining, hollowing-out machining, drilling, oblique holes, oblique cutting, etc. It is a means to solve the machining of impellers, blades, ship propellers, heavy-duty generator rotors, steam turbine rotors, large diesel engine crankshafts, etc.

What are the 5 axis? in a five-axis machining center?

X-axis: The X-axis is the horizontal axis of the machining center, controlling the movement of the tool in the horizontal direction. The movement of the X-axis affects the lateral position of the workpiece, determining its position and shape on the horizontal plane.

Y-axis: The Y-axis is the longitudinal axis of the machining center, controlling the movement of the tool in the longitudinal direction. The movement of the Y-axis affects the longitudinal position of the workpiece, determining its position and shape on the longitudinal plane.

Z-axis: The Z-axis is the vertical axis of the machining center, controlling the movement of the tool in the vertical direction. The movement of the Z-axis affects the height position of the workpiece, determining its position and shape on the vertical plane.

A-axis: The A-axis is the rotation axis of the four-axis and five-axis machining centers, controlling the rotation of the workpiece on the horizontal plane. The rotation of the A-axis allows the workpiece to rotate around the vertical axis on the horizontal plane, enabling multi-face machining.

C-axis: The C-axis is the rotation axis of the five-axis machining center, controlling the rotation of the workpiece around an axis perpendicular to the table. The rotation of the C-axis allows the workpiece to rotate around the rotation axis on the vertical plane, enabling more complex multi-angle and surface machining.

Types of Five-Axis Machining Centers

Vertical Five-Axis Machining Center

This type of machining center has two methods for the rotation axis: one is the worktable rotation axis.

The worktable, which is set on the bed, can rotate around the X-axis, defined as the A-axis. The general working range of the A-axis is +30 degrees to -120 degrees. There is also a rotating table in the middle of the worktable, which rotates around the Z-axis at the position shown in the illustration, defined as the C-axis. The C-axis can rotate 360 degrees. With the combination of the A-axis and C-axis, the workpiece fixed on the worktable can be machined on all five sides except the bottom surface by the vertical spindle. The minimum indexing value for the A-axis and C-axis is usually 0.001 degrees, which allows the workpiece to be subdivided into any angle for machining inclined surfaces and oblique holes.

If the A-axis and C-axis are linked with the XYZ three linear axes, complex spatial surfaces can be machined. Of course, this requires the support of high-end CNC systems, servo systems, and software. The advantage of this setup is that the structure of the spindle is relatively simple, the spindle rigidity is very good, and the manufacturing cost is relatively low.

Spindle Rotation Vertical Five-Axis Machining Center

The vertical machining center’s spindle has gravity acting downward, and the radial force on the bearings during high-speed idle operation is equal, resulting in good rotation characteristics. Therefore, the speed can be increased, with a general high speed reaching over 12,000 rpm, and the practical maximum speed has reached 40,000 rpm. The spindle system is equipped with a circulating cooling device, where the circulating cooling oil carries away the heat generated by high-speed rotation, is cooled to an appropriate temperature through a chiller, and then flows back into the spindle system.

The X, Y, Z three linear axes can also use linear encoders for feedback, with bidirectional positioning accuracy within the micron level. Since the rapid feed reaches 40-60 m/min or more, the ball screws for the X, Y, Z axes mostly adopt central cooling. Similar to the spindle system, the heat is carried away by the circulating oil that flows through the center of the ball screws after being chilled.

Summary

In the field of CNC machine tools, three-axis machining centers have relatively simple processing capabilities, four-axis machining centers can achieve multi-face machining, and five-axis machining centers have a higher level of multi-angle and surface machining capabilities. The main difference between three-axis, four-axis, and five-axis machining centers lies in the number of axes and processing capabilities. The choice of a suitable machining center should be determined based on specific processing requirements and complexity.