Введение

Water jet machining, also known as water jet cutting, is a versatile and innovative manufacturing process that uses a high-pressure stream of water to cut through various materials. This technology has gained significant popularity in industries ranging from aerospace to automotive, and from food processing to art and design. Water jet machining is renowned for its precision, flexibility, and environmental friendliness. Unlike traditional cutting methods that rely on heat or mechanical force, water jet cutting uses the kinetic energy of water to achieve clean, precise cuts without altering the material’s intrinsic properties. This article explores the principles, applications, advantages, and limitations of water jet machining, as well as its future potential in modern manufacturing.

The Principles of Water Jet Machining

Water jet machining operates on a simple yet powerful principle: a high-pressure stream of water is directed at a material to erode and cut through it. The process can be divided into two main types:

1. Pure Water Jet Cutting: This method uses only water, pressurized to levels as high as 60,000–90,000 psi (pounds per square inch). The water is forced through a small nozzle, typically made of sapphire or diamond, to create a fine, high-velocity stream. Pure water jet cutting is ideal for softer materials like rubber, foam, paper, and food products.
2. Abrasive Water Jet Cutting: For harder materials such as metals, ceramics, and composites, an abrasive substance (usually garnet) is added to the water stream. The abrasive particles accelerate the cutting process by enhancing the erosive power of the water jet. This method can cut through materials several inches thick with remarkable precision.

The key components of a water jet machining system include:

• High-Pressure Pump: Generates the ultra-high-pressure water stream.
• Nozzle: Focuses the water into a fine, high-velocity jet.
• Abrasive Delivery System: Introduces abrasive particles into the water stream (for abrasive cutting).
• CNC Controller: Guides the nozzle along the desired cutting path with high accuracy.
• Catcher Tank: Collects the spent water and abrasive particles after cutting.

Applications of Water Jet Machining

Water jet machining is used across a wide range of industries due to its versatility and precision. Some of the most notable applications include:

1. Metal Fabrication

Water jet cutting is widely used in the metalworking industry to cut materials such as steel, aluminum, titanium, and copper. Its ability to cut without generating heat makes it ideal for materials that are sensitive to thermal distortion. This is particularly important in aerospace and automotive industries, where precision and material integrity are critical.

2. Stone and Tile Cutting

In the construction and interior design industries, water jet cutting is used to shape natural stone, ceramic tiles, and glass. The process allows for intricate designs and precise cuts, making it a favorite for creating decorative elements and custom fixtures.

3. Food Processing

Water jet cutting is a hygienic and efficient method for cutting food products. It is used to slice bread, cut meat, and portion fish without compromising food safety or quality. The absence of heat ensures that the food’s texture and flavor remain intact.

4. Composites and Plastics

Water jet machining is ideal for cutting composite materials, which are often challenging to process using traditional methods. It is used in the production of carbon fiber components, fiberglass, and other advanced materials.

5. Art and Design

Artists and designers use water jet cutting to create intricate patterns and shapes in materials like wood, acrylic, and metal. The technology enables the production of highly detailed and customized pieces.

6. Medical Device Manufacturing

In the medical industry, water jet cutting is used to fabricate precision components for devices such as implants, surgical instruments, and diagnostic equipment. The process ensures clean edges and minimal material waste.

Advantages of Water Jet Machining

Water jet machining offers numerous advantages over traditional cutting methods, making it a preferred choice for many applications:

1. No Heat-Affected Zone (HAZ): Unlike laser or plasma cutting, water jet cutting does not generate heat, eliminating the risk of thermal distortion, warping, or changes in material properties.
2. Versatility: Water jet cutting can handle a wide range of materials, from soft and delicate substances to hard and durable ones. This makes it a one-stop solution for many industries.
3. Precision: The process allows for extremely tight tolerances, with cutting accuracy as high as ±0.001 inches. This level of precision is essential for industries like aerospace and medical device manufacturing.
4. Environmental Friendliness: Water jet cutting is a clean process that produces no harmful fumes, dust, or waste. The water used can often be recycled, and the abrasive materials are non-toxic.
5. Minimal Material Waste: The narrow kerf (cut width) of the water jet reduces material waste, making it a cost-effective option for expensive materials.
6. No Tool Wear: Since water jet cutting does not involve physical contact between a tool and the workpiece, there is no tool wear, reducing maintenance costs.
7. Ability to Cut Complex Shapes: The CNC-controlled nozzle can follow intricate paths, enabling the cutting of complex geometries and fine details.

Limitations of Water Jet Machining

Despite its many advantages, water jet machining does have some limitations:

1. Cutting Speed: While water jet cutting is precise, it can be slower than other methods like laser or plasma cutting, especially for thick materials.
2. Material Thickness: Although water jet cutting can handle thick materials, the process becomes less efficient as thickness increases. For extremely thick materials, alternative methods may be more suitable.
3. Operating Costs: The high-pressure pumps and abrasive materials can be expensive to maintain and replace, leading to higher operating costs compared to some traditional methods.
4. Surface Finish: While water jet cutting produces clean edges, the surface finish may require additional post-processing for certain applications.
5. Noise and Vibration: The process can generate significant noise and vibration, which may require mitigation measures in some environments.

Innovations and Future Trends

Water jet machining continues to evolve, driven by advancements in technology and the growing demand for precision manufacturing. Some of the key trends and innovations in the field include:

1. Hybrid Cutting Systems: Combining water jet cutting with other technologies, such as laser or plasma cutting, to leverage the strengths of each method.
2. Automation and Robotics: Integrating water jet cutting systems with robotic arms and advanced CNC controls to enhance precision and efficiency.
3. 3D Water Jet Cutting: Developing systems capable of cutting complex three-dimensional shapes, opening up new possibilities for manufacturing and design.
4. Eco-Friendly Abrasives: Research into alternative abrasive materials that are more sustainable and environmentally friendly.
5. Improved Pump Technology: Advances in high-pressure pump design to increase efficiency and reduce energy consumption.
6. AI and Machine Learning: Using artificial intelligence to optimize cutting parameters and improve process control.

Вывод

Water jet machining is a transformative technology that has revolutionized the way materials are cut and shaped. Its ability to deliver precision, versatility, and environmental benefits makes it an indispensable tool in modern manufacturing. While it has some limitations, ongoing advancements in technology are addressing these challenges and expanding the potential applications of water jet cutting. As industries continue to demand higher levels of precision and efficiency, water jet machining is poised to play an increasingly important role in shaping the future of manufacturing. Whether it’s cutting intricate designs in metal, slicing food products, or fabricating medical devices, water jet machining proves that sometimes, the simplest element—water—can be the most powerful tool.

Wear and deformation caused by the intense flow of the deformed metal on the die surface. Sometimes, the die cavity surface is scratched, abraded, or even adhered to, which easily causes dimensional errors and scrapping of the die, resulting in poor product quality.

The die cavity is easily deformed due to the relatively insufficient yield strength of the matrix material.

After continuous and repeated stress, especially in areas with stress concentration, the fatigue life of the alloy steel is relatively low, leading to fatigue failure as the number of heading operations increases.

Therefore, in addition to further improving the life of alloy steel dies, there is a widespread demand for a new material that is both wear-resistant and has good toughness to adapt to the development of the cold heading industry.

Test Results and Application Assessment

To expand the application field of cemented carbide and meet the urgent demand of steel ball factories for large-sized steel ball cemented carbide dies, based on our long-term work experience and domestic and foreign data collection, we finally selected a specially made coarse-grained WC powder, supplemented by an appropriate binder phase, to formulate YGC-I and YGC-Ⅱ coarse-grained cemented carbides (YGC-Ⅱ has a higher binder phase content than YGC-I). The conventional performance data of YGC-I and YGC-Ⅱ are shown in Table 1.

To study the effect of different particle sizes of WC powder on the fracture toughness of cemented carbide, fine-grained cemented carbide YG10X, medium-coarse grained cemented carbide YG18C, and YG20C were also prepared.

The fracture toughness of fine-grained, medium-coarse grained, and super-coarse grained cemented carbides with three different WC particle sizes was measured, and the results are shown in Table 2. It can be observed from Table 2 that as the WC particle size changes from fine-grained to medium-coarse grained and then to coarse grained, the fracture toughness of the alloy increases significantly, or even doubles. Even within the same particle size range, the fracture toughness of the alloy also increases with the increase of the binder phase content. This indicates that there are two ways to improve the fracture toughness of the alloy: first, by selecting coarse-grained WC powder, and second, by increasing the binder phase content of the alloy.

Selected YGC-I and YG10X samples were used to take scanning electron microscopy (SEM) fracture photos. Figure 1a shows the SEM fracture photo of YGC-I, and Figure 1b shows the SEM fracture photo of YG10X. Careful observation of the SEM fracture photos reveals that YGC-I has coarse grains and a thick binder phase layer, with cracks extending along the binder phase, resulting in mostly intergranular fractures and thus a higher fracture toughness K?C. In contrast, YG10X has very fine grains and a very thin binder phase layer, with mostly transgranular fractures, leading to a lower fracture toughness K?C. This indicates that the trend in fracture toughness values of carbides with different grain sizes is consistent with the conclusions drawn from the analysis of SEM fracture photos.

The aforementioned coarse-grained cemented carbide developed was applied to cold heading larger specification steel balls φ17.4625 and φ19.05. After being used on a Z32-28 machine at a certain factory, the usage data show that the coarse-grained cemented carbide is capable of stamping larger specification steel balls, and its die core life is more than 8 to 10 times longer than that of GCr15. This not only expands the application scope of cemented carbide but also provides the possibility for the bearing industry and standard parts industry to adopt carbide dies with larger heading forces. Moreover, due to the significant increase in die life, it brings about considerable economic benefits.

Вывод

1.Selecting specially made coarse-grained WC powder and correspondingly increasing the binder phase content of the alloy is an effective way to improve the fracture toughness of cemented carbide.

2.SEM fracture analysis indicates that coarse-grained cemented carbide fractures are mostly intergranular, resulting in a higher fracture toughness K?C, while fine-grained cemented carbide fractures are mostly transgranular, leading to a lower fracture toughness K?C.

3.The usage results show that the life of coarse-grained cemented carbide die cores is more than 8 to 10 times longer than that of GCr15.

4.In summary, for the stamping of larger specification steel balls, choosing coarse-grained cemented carbide is a reasonable option.

Functions of Cutting Fluid

Imagine the harsh working conditions when a twist drill rotates at astonishing speeds, penetrating deep into the workpiece for deep hole drilling. High temperatures, high pressures, and rapid rotation are extreme conditions that put immense stress on both the cutting tool and the workpiece. This is where cutting fluid steps in as a silent hero, playing a crucial role in cooling, lubricating, and cleaning during the drilling process.

Cooling Function

One of the primary functions of cutting fluid is cooling. During twist drill deep hole drilling, the friction between the tool and the workpiece generates a significant amount of heat. Without timely cooling, the tool is prone to damage due to overheating. Cutting fluid acts like a refreshing spring, carrying away the heat, protecting the tool, and ensuring the smooth progress of the drilling process.

Lubrication Function

In addition to cooling, cutting fluid also serves as a lubricant. During drilling, the contact area between the tool and the workpiece is very small, yet the pressure exerted is very high. Without sufficient lubrication, the tool can easily scratch the surface of the workpiece, affecting the quality of the hole. Cutting fluid acts like a lubricating film, reducing friction between the tool and the workpiece, decreasing wear, and improving the finish of the drilled hole.

Cleaning Function

Certainly, the cleaning function of cutting fluid should not be overlooked. During the drilling process, a significant amount of chips and metal powder is produced. If these chips are not removed in a timely manner, they can easily accumulate inside the drilled hole, leading to blockages and even damaging the tool. Cutting fluid acts like a diligent cleaner, continuously flushing the inside of the hole, carrying away the chips and metal powder, ensuring the hole remains unobstructed.

Distribution of Cutting Fluid

However, the distribution of cutting fluid in twist drill deep hole drilling is not uniform. Due to the limitations of the depth and diameter of the hole, it is difficult for the cutting fluid to reach the bottom of the hole directly. In some areas of the hole, the cutting fluid may form dead zones where the flow rate is very slow or almost non-existent. This results in the tool not receiving adequate cooling and lubrication in these areas, increasing the risk of tool wear and hole blockages.

How Cutting Fluid Affects Chip Removal

Chip removal is equally crucial in twist drill deep hole drilling. Chips are metal fragments produced during the drilling process, and if not removed promptly, they can easily accumulate inside the hole, forming blockages. Once a blockage occurs, it not only affects the quality of the hole but also puts tremendous pressure on the tool, which can lead to tool breakage. Therefore, timely chip removal is key to ensuring the smooth progress of the drilling process.

In twist drill deep hole drilling, the chip conveyance mechanism is relatively complex. Due to the limitations of the hole’s depth and diameter, chips cannot be easily expelled through the hole. They need to navigate through the tiny gap between the tool and the workpiece and then be carried out of the hole with the flow of cutting fluid. However, this process is fraught with challenges. The shape, size, and density of the chips all affect their conveyance efficiency. If the chips are too large or dense, they can easily form blockages inside the hole, leading to drilling failures.

Using 3D Simulation to Optimize Cutting Fluid Distribution for Improved Chip Removal

To optimize the distribution of cutting fluid and the removal of chips, scientists have conducted extensive research. They have utilized advanced 3D multiphysics simulation methods to conduct detailed simulations and analyses of the twist drill deep hole drilling process. These simulations not only reveal the flow characteristics of the cutting fluid inside the hole but also demonstrate the chip conveyance mechanism within the hole. Through these simulations, scientists have gained a deeper understanding of the reasons behind uneven cutting fluid distribution and inefficient chip removal, providing strong support for optimizing the drilling process.

In simulation studies, the coupled particle simulation (SPH-DEM) method has played a significant role. This method accurately simulates the movement and interaction of cutting fluid and chips inside the drilled hole. Through the SPH-DEM method, scientists can observe the flow trajectory of the cutting fluid within the hole, as well as the conveyance process of chips in the cutting fluid. These observations not only validate the accuracy of the simulation method but also provide an important basis for optimizing the distribution of cutting fluid and chip removal strategies.

In addition to coupled particle simulation, CFD (Computational Fluid Dynamics) simulation has also played a crucial role in the analysis of cutting fluid flow. CFD simulation can model the flow state of cutting fluid inside the drilled hole, including parameters such as flow velocity, pressure, and temperature. Through the analysis of these parameters, scientists can understand the distribution of cutting fluid within the hole, as well as the impact of different cutting fluid parameters on hole quality. These analytical results are of significant guiding significance for optimizing cutting fluid formulations and process parameters.

In the experimental validation phase, scientists designed a series of experiments to verify the accuracy of the simulation results. They selected different cutting parameters, types of cutting fluids, and concentrations for the experiments, and recorded data such as the quality of the drilled holes, the flow rate of the cutting fluid, and the chip removal situation. By comparing the experimental data with the simulation results, the scientists found a good consistency between the two. This not only verified the reliability of the simulation method but also provided strong support for optimizing cutting fluid distribution and chip removal strategies.

During the experimental process, the scientists also discovered some interesting phenomena. For example, under certain cutting parameters, although the flow rate of the cutting fluid was high, the quality of the drilled holes was not ideal. After analysis, they found that this was due to the formation of dead zones of the cutting fluid inside the hole, which resulted in insufficient cooling and lubrication for the tool in certain areas. To address this issue, they adjusted the injection angle and flow rate of the cutting fluid, successfully improving the distribution of the cutting fluid and enhancing the quality of the drilled holes.

In addition, scientists have also found that the chip removal is closely related to parameters such as the flow rate and viscosity of the cutting fluid. When the flow rate of the cutting fluid is too high, chips are easily carried away by the fluid; however, when the viscosity of the cutting fluid is too high, chips tend to form blockages inside the hole. Therefore, when optimizing the cutting fluid formulation, it is necessary to consider the flow rate, viscosity, and other parameters of the cutting fluid to ensure the smooth removal of chips.

Резюме

Through extensive experimentation and simulation studies, scientists have successfully optimized the cutting fluid distribution and chip removal strategies in twist drill deep hole drilling. They have found that by adjusting parameters such as the injection angle, flow rate, and viscosity of the cutting fluid, the distribution of the cutting fluid inside the hole can be significantly improved; at the same time, by optimizing the structure of the cutting tool and cutting parameters, the efficiency of chip removal can also be enhanced. These research findings not only improve the quality and efficiency of twist drill deep hole drilling but also provide valuable references for precision manufacturing in other fields.

The distribution of cutting fluid and chip removal in twist drill deep hole drilling is a complex and important process. Through scientific experimentation and simulation studies, we can gain a deeper understanding of the physical mechanisms and influencing factors in this process; by optimizing cutting fluid formulations and process parameters, we can significantly improve the quality and efficiency of drilling. In the future, with the continuous advancement of technology and the development of the manufacturing industry, it is believed that twist drill deep hole drilling technology will have an even broader development prospects and a wider range of applications.

At present, the regions in China that are most active in the recycling and regeneration of waste cemented carbide include Jinan City in Shandong Province, Qinghe County in Hebei Province, Mudanjiang City in Heilongjiang Province, and Zhuzhou City and Changsha City in Hunan Province.

Economic Benefits of Recycling Waste Carbide

Tungsten is the main component of cemented carbide, accounting for almost 50% of the total tungsten usage in its production, with China’s share being around 40%. Relevant data indicate that the demand for cemented carbide in various countries will rise significantly from the end of this century to the beginning of the next. It is estimated that by 2000, the demand could reach 40,000 tons, about 1.5 times the current production. However, tungsten is a rare element, with a crustal abundance of only 1×10?%, and the currently exploitable tungsten is only sufficient for 50 years.

Although China is a major producer of tungsten, both its reserves and recoverable quantities are showing a decreasing trend. Therefore, the rational utilization and recycling of tungsten resources should be placed on our agenda as an urgent issue to be seriously studied. Assuming that China’s recycling rate of waste alloy increases from 10% to 20%, it would mean an annual increase of several hundred tons of tungsten production. This would require the provision of several thousand tons of tungsten concentrate (containing 65% WO?) as raw material, equivalent to the tungsten content of 220,000 tons of raw ore (with a grade of 0.5% WO?). Thus, vigorously recycling cemented carbide is of great significance for the rational utilization and protection of existing tungsten resources.

Another major component of cemented carbide is cobalt. Due to the lack of cobalt resources in China, a large amount of cobalt needs to be imported annually to meet production needs. At the current level, recycling several hundred tons of cemented carbide in China each year could recover several tens of tons of cobalt, thereby saving a significant amount of foreign exchange for?our?country.

Processing Techniques of Recycling Waste Carbide

It is reported that there are currently about 30 different processing techniques used for the recycling of waste cemented carbide. Below is a brief introduction to several of the most commonly used and effective techniques in production.

Nitrate Fusion Method

This method involves melting waste cemented carbide together with nitrate at temperatures ranging from 900°C to 1200°C, resulting in the formation of soluble sodium tungstate. The reaction equation is as follows:

At this stage, the cooled melt is crushed and then leached with water to obtain a sodium tungstate solution and cobalt residue, which are then processed through normal procedures.

The advantages and disadvantages of this method are as follows: it has a large processing capacity and a wide range of applications, but it suffers from low recovery rates, high costs, poor working conditions, and significant pollution.

High-Temperature Oxidation Method

This method involves placing the cemented carbide in a temperature range of 700–950°C to oxidize it in air or oxygen. During this process, oxygen reacts with the alloy through the following chemical reaction:

The oxidized product is a brittle substance that, when treated with sodium hydroxide or a mixture of sodium hydroxide and sodium carbonate in a high-pressure leaching device, yields a sodium tungstate solution. The cobalt remaining in the residue is separated out according to conventional processes.

Phosphoric acid leaching method

Immerse the waste carbide?in a phosphoric acid solution and leach at a temperature of 50-60°C. Phosphoric acid reacts with cobalt in the carbide?to form cobalt phosphate, which enters the solution and separates from tungsten carbide. The advantages and disadvantages are: since phosphoric acid is a weak acid, the problem of equipment corrosion is easily solved, making it suitable for processing various waste carbides. However, the recovered tungsten carbide has a high oxygen content, and the subsequent process flow is long.

Zinc melting method

React waste carbides with zinc at a temperature close to 900°C. The cobalt in the alloy forms a zinc-cobalt low-melting-point alloy, causing the tungsten carbide in the waste alloy to lose the cobalt’s bonding effect and become loose. Then, vacuum distillation is used to evaporate and recover the metal zinc.

After the zinc melting process, the waste carbide?consists of layers of tungsten carbide and cobalt layers arranged in a multi-layered and interlocking pattern. It is a loose bulk material, which, after crushing, becomes a recycled carbide?mixture.

The advantages and disadvantages of the zinc melting method are: the process and equipment are relatively simple, the actual recovery rate is high, the production process causes less pollution, and the recovered mixture can be used directly for the production of tungsten products. However, this method consumes a lot of energy, with an electricity consumption of 4000-10000 degrees for processing 1 ton of waste carbide, and the recovered material contains a small amount of zinc, which has a certain impact on product quality.

Sodium sulfate fusion method

This method involves reacting waste carbides with sodium sulfate at a temperature of 900-1000°C to form a molten tungstic acid. After cooling, it is then leached with hot water to obtain a sodium tungstate solution and cobalt slag. The reaction equation is as follows: (The specific reaction equation is not provided in the original text, so it cannot be translated here.)

Its advantages and disadvantages are: wide adaptability and large production capacity. The drawback is that sulfur dioxide gas is emitted during the production process.

Electrolysis fusion method of recycling waste carbide

This method involves placing the waste carbide?in an electrolytic cell as the anode, nickel plate as the cathode, and dilute hydrochloric acid as the electrolyte. After electrolysis, the cobalt in the carbide?enters the solution in the form of COCl?. The washed and ground WC can be directly used to produce alloys. This method yields pure products, is highly efficient, has simple equipment, and is easy to operate. It is particularly suitable for processing with high cobalt content and has high value for promotion and application.

Cold embrittlement method

The cold embrittlement method involves crushing the waste carbide?coarsely, removing impurities, and then using a high-speed air stream to inject the coarsely crushed carbide?into a vacuum chamber equipped with a carbide?paddle, followed by further crushing to obtain a mixture.

This method has a wide range of processing and treatment, and the production process does not cause environmental pollution, but the equipment cost is relatively high.

According to reports, the zinc dissolution method is widely used for the recycling of waste carbides in China at the current stage. Overseas, the most economical method is considered to be a combination of the zinc dissolution method and the cold embrittlement method. In summary, there are many methods for recycling and processing waste carbides, each with its own pros and cons. When selecting a method in practice, a comprehensive analysis and comparison should be conducted based on the type of waste carbide, the size of the production scale, equipment capacity, technical level, and the source of raw and auxiliary materials, to choose an advanced, reasonable, and economically significant process for production practice.

Technical and Economic Preliminary Evaluation of Several Process Methods for Recycling Waste Carbides

As mentioned earlier, the zinc fusion method, electrolysis dissolution method, and mechanical crushing method have all become the main industrial methods for recycling and regenerating waste carbides. The recycled and regenerated powders can be used to produce carbides through conventional processes, which undoubtedly promotes the full utilization of non-ferrous metal resources such as tungsten and cobalt, saves energy, reduces manufacturing costs, promotes the development of small and medium-sized enterprises, and provides employment for the unemployed. Among them, the powder produced by the electrolysis dissolution method is particularly good in quality with low impurity content, low energy consumption, moderate processing costs, and good economic benefits (the selling price of WC powder is only 60% to 70% of the conventional product). It is one of the main methods being vigorously developed in many regions of our country.

Резюме

It should be pointed out that the recycling and regeneration of waste carbides is a new venture and an emerging industrial sector in the field of carbides, which should be affirmed and supported. At the same time, great attention must be paid to product quality in the work of recycling and utilizing waste carbides, especially since most enterprises are still at a relatively low level of manual workshop production. Many enterprises often focus only on economic benefits while neglecting the control and testing of the recycling process and powder quality, which is incorrect. In the future, enterprises engaged in this kind of production should continuously improve and enhance the recycling process and product quality, strengthen the recognition of the importance of “product quality is the life of the enterprise,” and must not be careless. Otherwise, it will be difficult to maintain a firm foothold in the fierce market competition for a long time, let alone continue to develop and grow.

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.

Вывод

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 карбид 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.

Вывод

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 Режущий инструмент

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.

Вывод

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.

Резюме

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 карбид 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.

Вывод

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