The present invention relates to a rotary cutting tool comprising a maraging steel carrier body and a cutting element wherein the carrier body and the cutting element are joined by brazing. The present invention also relates to a method of making such rotary cutting tool.

Background

It is common in the art to make rotary cutting tools such as drills or endmills for chip forming machining from a round tool blank of cemented carbide. The shape of the drill or endmill, e.g. chip flutes, cutting edges etc. is achieved by grinding. Grinding is time consuming and also expensive. Due to this, rotary cutting tools can be reconditioned several times by regrinding the worn parts like the cutting edge and also be provided with a new coating.

There have been several attempts to replace part of the rotary cutting tool body, e.g. the shank portion with another, less costly, material. It is also known to have a loose, replaceable top on a drill that is fastened by mechanical means, like e.g. with a screw.

Welding or brazing a cutting element to a carrier body of a different material than in the cutting element is known in the art.

Using steel as a carrier material is usually not considered to be an option due to problems with brazing and low hardness and/or tensile strength of the steel. Since a cutting tool is subjected to large forces when used in cutting operations, the braze joint needs to be strong and the carrier body needs to have an optimum toughness/hardness ratio. Also, the difference in CTE (coefficient of thermal expansion) between steel and materials such as cemented carbide, polycrystalline diamond (PCD) or a cubic boron nitride (cBN) can lead to cracks.

Also, for rotary cutting tools, which are usually provided with cooling channels, joining a cutting element with a carrier body by brazing could lead to clogging of the cooling channels with braze material.

Although cemented carbide seems suitable to be used as a carrier body, it still has its disadvantages. For environmental reasons, recycling of cemented carbide is preferred which is a complicated process. Also, cemented carbide is hard to machine and arriving at the final shape of a rotary cutting tool usually requires extensive grinding etc.

Although maraging steel works well with regard to strength of the braze joint etc. it can still have its disadvantages as a carrier body in certain cutting applications since maraging steel will not reach the same level of hardness as cemented carbide. The lower hardness will decrease the wear resistance. For example, in drilling where the work piece chip will hit the carrier body and not just the cutting element, the wear resistance of the maraging steel might not be good enough.

One object of the present invention is to provide a rotary cutting tool which has a steel carrier body that can withstand the forces during metal cutting operations.

Another object of the present invention is to provide a rotary cutting tool with a cutting element joined to the steel carrier body with a high strength braze joint.

Another object of the present invention is to provide a rotary cutting tool where the carrier body can be shaped with less effort as compared to a cemented carbide carrier body.

Another object of the present invention is to provide a rotary cutting tool where the carrier body can withstand a high degree of wear from work piece chips.

Definitions

By rotary cutting tool is herein meant any round tool used for chip forming machining cutting operations. Examples are drills, endmills and reamers.

A rotary cutting tool (also called solid round tool) typically comprises a shank portion and a fluted portion integral with the shank portion. The fluted portion refers to the portion of the rotary cutting tool having chip flutes formed in a circumferential surface thereof. One or more interior coolant channels may be provided within the rotary cutting tool.

By cutting element is herein meant the part of the cutting tool insert that is engaged in the cutting operation, i.e. the part that comprises at least one cutting edge and is in contact with the work piece. The cutting element can have different shapes depending on the cutting application. As an example, the cutting element can constitute the whole top part of the rotary cutting tool, see Figure 1, or the cutting element can be smaller, like e.g. a vein or nib, that is brazed to the outermost part of the rotary cutting tool, see Figure 2.

By carrier body is herein meant the cutting tool body, that does not constitute the cutting element. The carrier body can have any shape of a rotary cutting tool see above, and preferably comprises at least part of the fluted portion.

Detailed description of the invention

The present invention relates to a rotary cutting tool comprising a carrier body, and at least one cutting element comprising at least one cutting edge, a braze joint, joining said carrier body and said at least one cutting element, where said braze joint comprises Ti and wherein said braze joint comprises a Ti containing joining layer with a thickness of between 0.03 and 5 pm adjoining to the cutting element The carrier body is made of maraging steel.

The cutting element can be made of any material known in the art of metal cutting, i.e. one of cemented carbide, cermets, ceramics, Polycrystalline diamond (PCD), or sintered cubic boron nitride (PcBN).

By ceramic is herein meant a material comprising transition metal carbides, nitrides or carbonitrides grains e.g. WC, Si ₃ N ₄ , SiAION, AI ₂ O ₃ , SiC whiskers etc. embedded in an oxide ceramic matrix e.g. aluminum oxide, where the amount of transition metal carbides, nitrides or carbonitrides grains is between 5 to 45 vol%. They are generally sintered in a hot isostatic pressing process.

The cemented carbide used as a cutting element can be made of any cemented carbide known in the art. The cemented carbide comprises a hard phase embedded in a metallic binder phase matrix.

By cemented carbide is herein meant that at least 50 wt% of the hard phase is WC.

Suitably, the amount of metallic binder phase is between 3 and 20 wt%, preferably between 4 and 15 wt% of the cemented carbide. Preferably, the main component of the metallic binder phase is selected from one or more of Co, Ni and Fe, more preferably the main component of the metallic binder phase is Co. By main component is herein meant that no other elements other than those mentioned above are added to form the binder phase. However, if other components are added, like e.g. Cr, it will inevitably be dissolved in the binder during sintering.

In one embodiment of the present invention, the cemented carbide can also comprise other components common in cemented carbides elements selected from Cr, Ta, Ti, Nb and V present as elements or as carbides, nitrides or carbonitrides.

By cermet is herein meant a material comprising hard constituents in a metallic binder phase, wherein the hard constituents comprise carbides or carbonitrides of one or more of Ta, Ti, Nb, Cr, Hf, V, Mo and Zr, such as TiN, TiC and/or TiCN.

By PCD (polycrystalline diamond) is herein meant a material comprising diamond crystals sintered together where the amount of diamond crystals is between 50 to 100 vol%. The diamond crystals typically have a grain size of between 0.5 and 30 pm. The PCD can also comprise one or more constituents selected from Al, Cr, Co, Ni, V, Fe and Si.

By PcBN is herein meant a material comprising cBN grains embedded in a metallic and/or ceramic binder where the amount of cBN grains is between 30 to 99 vol%. The ceramic binder can contain one or more constituents being carbides, nitrides, carbonitrides, borides or oxides of elements selected from Co, Ni and groups 4-6 in the periodic table of elements.

Polycrystalline diamond (PCD) and sintered cubic boron nitride (PcBN) can either be provided as it is, so called "free standing" or together with a cemented carbide support, so called "carbide backed". Polycrystalline diamond (PCD) and sintered cubic boron nitride (PcBN) are usually manufactured by providing a suitable powder mixture which is subjected to a high temperature-high pressure (HP/HT) sintering step to form a sintered compact (typically 1400°C, 5GPa).

When a polycrystalline diamond (PCD) and sintered cubic boron nitride (PcBN) are provided with a cemented carbide support, this is prepared already prior to the sintering of the polycrystalline diamond (PCD) and sintered cubic boron nitride (PcBN). One way of doing this is to use a cup with a cemented carbide disc in the bottom. The cup is then filled with the PCD or cBN powder mixture of choice and the cup is then sealed. The sealed cup is then subjected to a high temperature-high pressure (HPHT) sintering step. The diamond or cBN material is bonded to the cemented carbide during the sintering step. The disc can then be cut into suitable pieces using e.g. laser or WEDM (wire electrical discharge machining).

The cemented carbide used as a support for the polycrystalline diamond (PCD) and sintered cubic boron (PcBN) can be made of any cemented carbide common in the art, see the definition above.

Maraging steels suitably contain from 8 to 25 wt% Ni and one or more alloying elements selected from Co, Mo, Ti, Al and Cr in a total amount of between 7 and 27 wt%, preferably between 7 and 23 wt% of alloying elements. Maraging steels typically contain less carbon than conventional steel, suitably 0.03 wt% or less. The balance being Fe.

In one embodiment of the present invention the maraging steel contains from 11 to 25 wt% Ni, preferably 15 to 25 wt% Ni. The alloying elements are suitably Co in an amount of from 7 to 15 wt%, preferably 8.5 to 12.5 wt% Co, Mo in an amount of from 3 to 10 wt%, preferably 3 to 6 wt% Mo, Ti in an amount of from 0.1 to 1.6 wt% preferably from 0.5 to 1.2 wt% Ti, from 0 to 0.15 wt% Cr, Al in an amount of from 0 to 0.2 wt% and less than 0.03 wt% C. The balance being Fe.

In one embodiment of the present invention, the maraging steel has a composition of from 17 to 19 wt% Ni, from 8.5 to 12.5 wt% Co, from 4 to 6 wt% Mo, from 0.5 to 1.2 wt% Ti, from 0 to 0.15 wt% Cr, from 0 to 0.2 wt% Al and less than 0.03 wt% C. The balance being Fe.

In another embodiment of the present invention, the maraging steel has a composition of from 8 to 11 wt% Ni, preferably from 9 to 10 wt% Ni, from 2.5 to 4 wt% Cr, preferably from 3 to 3.5 wt% Cr, from 3.5 to 5 wt% Mo preferably from 4 to 4.5 wt% Mo, from 0.4 to 1.1 wt% Ti, preferably from 0.7 to 0.9 wt% Ti, less than 0.4 wt% Si, less than 0.4 wt% Mn and balance Fe.

As all alloys, maraging steel can also contain unavoidable impurities. By impurities is herein meant any element that can be present in the maraging steel in such small amounts that it does not have any influence on the properties of the steel. The total amount of impurities is below 0.50 wt%, preferably below 0.15 wt%. Examples of such elements are Mn, P, Si, B and S. In one embodiment of the present invention, the amount of Mn is less than

0.05wt%, the amount of P is less than 0.003 wt%, the amount of Si is less than 0.004 wt% and S less than 0.002 wt%.

The average hardness of the maraging steel part will depend on if any ageing/nitriding step has been performed or not, see below.

In one embodiment of the present invention the carrier body made of maraging steel is not provided with a gradient with regard to hardness the average hardness is between 300 and 1200 HV1, preferably between 500 and 1100 HV1. The standard deviation of the hardness values is suitably between 0 to 150 HV1, preferably between 0 and 100 HV1.

In one embodiment of the present invention, the carrier body made of maraging steel is provided with a hardness gradient, i.e. the carrier body has an increased hardness in a surface zone compared to in the core. By that is herein meant that the hardness has its highest value at the surface and then gradually decreases towards the core. The carrier body made of maraging steel then has an average core hardness of between 300 and 700 HV1, preferably between 500 and 700 HV1. The standard deviation of the core hardness values is suitably between 0 to 20 HV1, preferably between 0 and 15 HV1. The surface of maraging steel then has an average surface hardness of between 300 and 1200 HV1, preferably between 500 and 1100 HV1. The standard deviation of the hardness values is suitably between 0 to 150 HV1, preferably between 0 and 100 HV1. The surface hardness is at least 30 % higher than the core hardness, preferably at least 40 % higher than the core hardness.

By "core" is herein meant the inner part of the maraging steel carrier body, where the hardness, when measured on a cross section, is no longer changing.

The depth of the hardness gradient as measured from the surface, the nitriding depth, is determined by making a hardness depth curve on the transverse section of a maraging steel carrier provided with a hardness gradient, measuring HV 0.3 or HV0.5 according to the standard DIN EN ISO 6507-1, starting close to the surface and towards the core until the hardness is no longer changing. The nitriding depth is given by the vertical distance from the surface of the nitrided carrier body up to the point of the limiting hardness where the limiting hardness is defined as the average core hardness + 50 HV0.3 or 50 HV0.5, see figure 4. The average nitriding hardness depth of the maraging steel carrier body is between 0.001 and 0.8 mm, preferably between 0.01 and 0.3 mm. The standard deviation of the hardness values is suitably between 0 to 0.03 mm preferably between 0 and 0.02 mm.

By increasing the hardness on the surface of the maraging steel carrier, it will have an increased wear resistance. This can be a big advantage when the cutting tool insert according to the present invention is used in the cutting applications where the chip from work piece material hits the maraging steel carrier.

The brazing technique is the so-called active brazing. By that is meant that the joint is not just formed by melting the filler material and forming a metallic bond, it also involves a chemical reaction with one or both of the materials that are to be joined. The joining element in the filler material is usually Ti, however elements such as Hf, V, Zr and Cr are also considered to be active elements. According to this invention, Ti is the active element.

By braze joint is herein meant the area or mass between the cemented carbide and the maraging steel part that is filled by the filler material and formed during the brazing process, see below.

The thickness of the braze joint is suitably between 5 and 200 pm, preferably between 15 and 100 pm.

The braze joint is not a homogenous phase. Instead, after brazing, the elements in the filler material form different alloying phases.

The braze joint, after brazing, comprises a Ti containing joining layer adjoining to the cutting element. Ti is very reactive and will, during brazing, react with one or more elements present in the cutting element. Most commonly, covalent bonds are formed with one or more of carbon, nitrogen, oxygen and boron and form a strong Ti containing joining layer at the interface between the braze joint and the cutting element.

The composition of the Ti containing joining layer will vary depending on what material the cutting element is made of but is usually composed of one of TiC, TiN, TiO _x and TiB _x or a mixture thereof. Since the formed joining layer is of ceramic nature, the joint may become brittle if the layer growth is uncontrolled. For example, if the material closest to the braze joint is PCD (polycrystalline diamond) or cemented carbide, either that the whole cutting element is made of cemented carbide or if it is a carbide backed PCD or PcBN cutting element, the Ti containing joining layer is a TiC layer. The Ti in the braze joint will react with the carbon in the WC or diamond and form TiC.

Another example is that if the cutting element is made of solid (also called "free standing") PcBN, the joining layer will be TiN since Ti will react with the nitrogen in the cBN, but can also contain smaller amounts of TiB _x , like e.g. TiB2.

When the cutting element is made of ceramics, e.g. a AI2O3/WC sintered ceramic composite the joining layer will be a TiC/TiOx layer.

There are several ways to detect the presence of a joining layer depending on which type of equipment that is used.

If a Scanning Electron Microscope (SEM) with a high enough resolution is used, the joining layer is clearly visible adjacent the cutting element. To verify the composition of the layer, SEM-EDS (energy dispersive spectroscopy) and/or SEM-EPMA (electron probe microscopy analysis) with WDS (wave length dispersive spectroscopy) can be used to identify the individual elements in the joining layer.

In one embodiment of the present invention, the thickness of the joining layer is between 0.03 and 5 pm, preferably between 0.05 and 1 pm, more preferably between 0.05 and 0.5 pm and most preferred between 0.05 and 0.25 pm.

If the SEM image used does not have enough resolution to detect the joining layer, the accumulation of Ti and/or C at the interface between the filler material and the cutting element can be seen using e.g. SEM-EDS or SEM-EPMA with WDS. The accumulation of Ti is herein after called the Ti-accumulation layer and is one indicator that a joining layer is formed, even if not visually detected in the SEM image. The Ti-accumulation layer is considerably thicker than the actual joining layer which could mean that not all Ti will form TiC/TiN/TiOx/TiBx. The thickness of the Ti-accumulation layer is also partly affected by the analysis method. Preferably, the braze joint, in addition to Ti, further comprises one or more elements selected from Ag, Cu, Sn, In, Zr, Hf and C, more preferably from Ag, Cu and In.

The braze joint can also contain smaller amounts of other elements considered to be unavoidable impurities. By unavoidable impurities are herein meant small amounts of elements possibly present in the braze material, other than those listed above, prior to the brazing step as well as elements from the materials to be joined, e.g. Co, W etc. from the cemented carbide and Fe, Ni etc. from the maraging steel. Small amounts of the elements from the parts to be joined are unavoidably dissolved in the braze material when subjected to the increased temperature during the brazing step whereby the braze material melts and allow diffusion from the joining parts. As long as the brazing process parameters such as temperature and time are within the ranges according to the present invention, the total amount of the unavoidable impurities will be so small that it does not affect the performance of the braze joint.

The composition of the braze joint after brazing is difficult to determine since the elements are not evenly distributed. If available, the easiest way is to look at the filler material that has been used since the paste or foil is a homogenous blend. Also, the braze joint might comprise small amounts of elements from the materials to be joined, e.g. Co, W from the cemented carbide and Fe, Ni etc. from the maraging steel.

The amount of Ti and possible further elements in the braze joint could also be measured using Energy-dispersive X-ray spectroscopy analysis (EDS). However, due to the uneven distribution of the precipitated elements in the braze joint, many measuring points need to be used and the standard deviation will be large. Preferably, the braze joint comprises, in average, Ag in an amount of from 30 to 80 wt%, preferably from 40 to 75 wt%, Cu in an amount of 15 to 50 wt%, preferably from 15 to 40 wt%, Ti in an amount of 0.3 to 15 wt%, preferably from 0.5 to 5 wt%, Sn in an amount of 0 to 10 wt%, preferably from 0 to 2 wt% and In in an amount of from 0 to 30 wt%, preferably from 5 to 25 wt%.

At the interface between the braze joint and the maraging steel part Ti is also accumulated in the braze joint where it forms a metallic bond with the iron in the steel. The thickness of the accumulation layer of Ti at the maraging steel surface is preferably between 1 and 10 pm, preferably between 2 to 5 pm and can be measured by e.g. EDS. In one embodiment of the present invention, the cutting element constitutes the top of the rotary cutting tool, also called head. By that is herein meant that the cutting element comprises the at least one cutting edge as well as a part of the fluted portion. If the rotary cutting tool is provided with one or more cooling channels, the cooling channels go through both the carrier body and the cutting element. One example of such a rotary cutting tool can be seen in Figure 1.

In one embodiment of the present invention, the cutting element is a nip or a vein. One example of such a rotary cutting tool can be seen in Figure 2.

The present invention also relates to a method of making a cutting tool according to the above comprising the steps of:

-providing a carrier body made of maraging steel,

- providing at least one cutting element comprising at least one cutting edge; -placing a filler material comprising Ti in an amount of 0.3 to 15 wt% of the filler material between and in contact with the carrier body and the cutting element;

-subjecting the carrier body and the cutting element with the filler material in between to a brazing step in a furnace at a temperature between 600 and 830°C, for a time period of between 1 and 60 minutes and wherein the brazing takes place in vacuum.

The filler material (also called braze metal) according to the present invention contains Ti in a total amount of from 0.3 to 15 wt%, preferably 1 to 5 wt% of the filler material. The filler material of the present invention suitably has a solidus temperature of between 490 and 1125°C, preferably between 600 and 700°C. Further, the filler material of the present invention has a liquidus temperature of between 610 and 1180°C, preferably between 700 and 750°C. The filler material further comprises, in addition to Ti, one or more elements selected from Ag, Cu, Sn, In, Zr, Hf and Cr.

In one embodiment of the present invention, the filler material comprises Ag in an amount of from 30 to 80 wt%, preferably from 40 to 75 wt%, Cu in an amount of 15 to 50 wt%, preferably from 15 to 40 wt%, Ti in an amount of 0.3 to 15 wt%, preferably from 0.5 to 5 wt%, Sn in an amount of 0 to 10 wt%, preferably from 0 to 2 wt% and In in an amount of from 0 to 30 wt%, preferably from 5 to 25 wt%.

Suitably, the filler material is provided as a foil or paste. The filler material is provided onto the joining surfaces of the cutting element and the maraging steel carrier.

The thickness of the filler material provided at the joining surfaces prior to the brazing process depends on the type of material, i.e. foil or paste. If a paste is used, enough material is applied so that the surface that is to be brazed is covered. Typically, the thickness is between 5 and 200 pm, preferably between 15 and 100 pm.

The parts, i.e. the cutting element and the maraging steel carrier with filler material in between, are then placed in a furnace with an inert or reducing environment, i.e. with a minimum amount of oxygen. Preferably, the brazing temperature in the furnace is between 600 and 830°C, preferably between 650 and 820°C more preferably between 700 and 750°C.

The choice of brazing temperature depends on several things. For example, if the cutting element is made of PCD, the brazing temperature should be below 750 °C to avoid graphitization of the diamonds, whereas if the cutting element is made of cemented carbide, the widest temperature range described above could be used.

The time the parts are subjected to the elevated temperature is between 1 and 60 minutes, preferably between 5 and 15 minutes. If the time at elevated temperature is shorter, there is not enough time for the braze joint to form and the Ti to react to reach the desired strength of braze joint. If the time at elevated temperature is longer, the Ti- containing, brittle reaction zone will grow uncontrollably, which negatively influences the joint properties, e.g. shear strength.

The brazing suitably takes place in vacuum or with the presence of Argon at low partial pressure. By vacuum is herein meant that the pressure in the furnace is below 5xl0 ^-4 mbar, preferably below 5xl0 ^-5 mbar. If argon is present, the argon pressure is below lxlO ^-2 mbar.

The brazing furnace used according to the present invention can be any furnace that can provide such well controlled conditions with regard to a vacuumizing, heating and cooling rate etc. as has been described above.

In one embodiment of the present invention the parts are subjected to an ageing step after the brazing step by subjecting the brazed parts to an elevated temperature of between 300 and 600°C, preferably between 350 and 500°C and most preferably between

400 and 440°C, for a time of between 5 minutes and 12 hours, preferably between 30 minutes and 8 hours and more preferably between 3 and 6 hours.

Suitably the heating rate up to the ageing temperature is between 1 to 50 °C/min, preferably between 5 to 10°C/min. Suitably the cooling rate from the ageing temperature down to a temperature of at least below the solidus temperature of the filler material, preferably below 300°C, is between 1 to 50°C/min, preferably between 5 to 10°C/min.

In one embodiment of the present invention, the ageing takes place directly after the brazing step in the same furnace as the brazing step takes place.

In one embodiment of the present invention, the ageing takes place directly after the brazing step in a different furnace from the vacuum brazing.

In one embodiment of the present invention, the ageing takes place in the same furnace/deposition chamber before or during deposition of a coating.

The ageing step will increase the overall hardness of the maraging steel.

In one embodiment of the present invention the ageing step is at least partly performed in a nitriding atmosphere. Due to the temperature during the nitriding, the ageing effect will also be there and therefore there is usually no additional separate ageing step if nitriding is performed.

The nitriding step can be performed using plasma nitriding or gas nitriding, preferably plasma nitriding. The nitriding atmosphere can be provided by a nitrogen containing gas e.g. N2, NH3.

In one embodiment of the present invention the nitriding step is performed using plasma nitriding. By that is herein meant that the nitriding takes place in vacuum vessel that is provided with a plasma generator where a nitriding atmosphere can be provided. The temperature is suitably between 300 and 600°C, preferably between 350 and 550°C and the duration can be between 1 to 100 hours. The pressure should preferably be low, suitably between 50 and 600 Pa. For plasma nitriding the gas is preferably N2, which can be mixed with e.g. H ₂ . In one embodiment of the present invention the nitriding step is performed using gas nitriding. The gas nitriding is preferably performed at a temperature between 450 and 600°C, preferably between 500 and 520°C. The gas nitriding is preferably done by NH3 which is split into H2 and N2 in the reactor. Gas nitriding can either be performed at low pressure, preferably 0.05-0.02 MPa, or close to atmospheric pressure.

The exact temperature, duration and choice of nitriding gas is dependent on several things, the desired nitriding effect on the maraging steel, the specific type of equipment that is used etc.

In one embodiment of the present invention, the maraging steel part has the following composition of from 18 to 19 wt% Ni, from 8 to 10 wt% Co, from 4 to 6 wt% Mo, from 0.5 to 1.2 wt% Ti, from 0 to 0.15 wt% Cr, from 0 to 0.2 wt% Al, less than 0.03 wt% C, less than 0.04 wt% Si, less than 0.05wt% Mn, less than 0.003 wt% P, less than 0.002 wt% S and less than 0.0005 wt% B. The balance being Fe. The filler material preferably has the following composition Ag in an amount of from 40 to 75 wt%, Cu in an amount of from 20 to 40 wt%, Ti in an amount of from 0.5 to 5 wt%, Sn in an amount of from 0 to 2 wt% and In in an amount of from 10 to 25 wt%.

In one embodiment of the present invention, the maraging steel part has the following composition of from 9 to 10 wt% Ni, from 3 to 3.5 wt% Cr, from 4 to 4.5 wt% Mo, from 0.7 to 0.9 wt% Ti, less than 0.4 wt% Si, less than 0.4 wt% Mn and balance Fe. The filler material preferably has the following composition Ag in an amount of from 40 to 75 wt%, Cu in an amount of from 20 to 40 wt%, Ti in an amount of from 0.5 to 5 wt%, Sn in an amount of from 0 to 2 wt% and In in an amount of from 10 to 25 wt%.

The most common way to make a rotary cutting tool made solely from one material, e.g. cemented carbide, is to start with a cylindrical rod which is grinded into its final shape i.e. with at least one flute and at least one cutting edge. When making a rotary cutting tool according to the present invention the brazing step can take place either before or after the formation of the flutes and at least one cutting edge. If the rotary cutting tool is subjected to an ageing step, either with or without nitriding, the formation of the flutes and at least one cutting edge preferably takes place after the brazing step but before the ageing step, either with or without nitriding. In one embodiment of the present invention, the at least one flute and the at least one cutting edge can be formed prior to the brazing step.

In one embodiment of the present invention, the rotary cutting tool can be coated with a wear resistant coating to further enhance the cutting performance. The coating can be deposited using PVD or CVD technique, however PVD is the most common technique for rotary cutting tools.

Drawings

Figure 1 shows an exemplary drill where the cutting element A is a drill head and B is the carrier body of maraging steel with a braze joint C in between. The drill has a shank portion 1, a fluted portion 2 with a flute 4 and also cooling channels 3.

Figure 2 shows a drill where the cutting element is a vein (can also be called nib) where cutting element A is a vein and B is the carrier body of maraging steel with a braze joint C in between (not visible).

Figure 3 shows a schematic drawing of the shear testing device where 1 is the steel part and 2 is the cemented carbide part and where F is the applied force.

Figure 4 shows an example of a hardness depth curve showing the hardness values decreasing from the surface towards the core where A is the average core hardness, B is the limiting hardness and C is the nitriding depth.

Figure 5 shows an SEM image of the braze joint where A is the cemented carbide, B is the braze joint, C is the maraging steel carrier and D shows the interface where the Ti containing layer is situated.

Example 1 (Invention)

Steel parts made of maraging steel 1.2709 in the form of a cylinder were provided together with cemented carbide parts with a composition of 10 wt% Co, 1 wt% other carbides and the remaining WC. The maraging steel had a hardness of approx. 340 HV1 prior to brazing.

The braze material (Incusil ABA from WBC Group) was provided in the form of a foil with a thickness of 100 pm. The braze material had a composition of 59.0 wt% Ag, 27.5 wt% Cu, 12.5 wt% In, and 1.25 wt% Ti. The solidus temperature was ca. 605°C, the liquidus temperature was ca. 715°C.

The foil was placed between the maraging steel part and the cemented carbide part so that both pieces were in contact with the foil. The assembled joining pieces were then placed into a Schmetz vacuum furnace (type: EU 80/1H 30 x 45 x 30 6 bar System *2RV*) where the temperature was first increased to 740°C at a rate of 20°C/min. The brazing temperature 740°C was kept for 15 minutes after which the pieces were cooled down to 300°C at a rate of 5°C/min. After 300°C it was free cooling to room temperature.

Excellent wetting with no signs of thermal stress crack could be observed, proven by the high shear test result.

This sample is herein denoted Invention 1.

Example 2 (Invention)

Steel parts made of maraging steel 1.2709 in the form of a cylinder were provided together with cemented carbide parts with a composition of 10 wt% Co, 1 wt% other carbides and the remaining WC. The maraging steel had a hardness of approx. 340 HV1 prior to brazing.

The braze material (TB-651 from Tokyo Braze) was provided in the form of a foil with a thickness of 100 pm. The braze material had a composition of 65.0 wt% Ag, 28.0 wt% Cu, 2.0 wt% Ti, and 5.0 wt% Sn. The solidus temperature was ca. 700°C, the liquidus temperature was ca. 750°C.

The foil was placed between the maraging steel part and the cemented carbide part so that both pieces were in contact with the foil. The assembled joining pieces were then placed into a Schmetz vacuum furnace (type: EU 80/1H 30 x 45 x 30 6 bar System *2RV*) where the temperature was first increased to 815°C at a rate of 20°C/min. The brazing temperature 815°C was kept for a time 15 minutes after which the pieces were cooled down to 300°C at a rate of 5°C/min. After 300°C it was free cooling.

An excellent wetting with no signs of thermal stress crack could be observed, proven by the high shear test result. This sample is herein denoted Invention 2.

Example 3 (Ageing)

After the brazing step, samples of Invention 1 and Invention 2 were subjected to an ageing process to retain the hardness of the maraging steel. The pieces were placed into a furnace where the temperature was first increased to 490°C at a rate of 5°C/min. The temperature 580°C was kept for 3h after which the pieces were cooled down to 300°C at a rate of 5°C/min. After 300°C it was free cooling. The results can be seen in Table 1.

Example 4 (Plasma nitriding)

Samples according to Invention 1 and 2 were subjected to a plasma nitriding step in a gas flow of 350:50 ml/min H2:N2 at a chamber pressure of 3 mbar. The temperature in the chamber was 490°C. The time the samples were subjected to the plasma nitriding step was 16 hours. No masking of the braze joint was used before the nitriding. A SEM image of a cross section of Invention 2 plasma nitrided for 16h is shown in figure 5.

Example 5 (gas nitriding)

A sample according to Invention 1 was subjected to a gas nitriding step by NH ₃ splitting. The temperature in the chamber was 510°C. The time the samples were subjected to the gas nitriding step was 23 or 55 hours. No masking of the braze joint was used before the nitriding.

Example 6 (comparative)

A steel part made of the carbon-hardening cold-work steel 1.2714 was provided together with a cemented carbide part with a composition of 10 wt% Co, 1 wt% other carbides and the remaining WC.

The braze metal was provided in the form of a foil with a thickness of 100 pm. The braze material 1 had a composition of 100.0 wt% Cu. The melting temperature is 1085 °C.

The foil was placed between the steel part and the cemented carbide part and assembled joining pieces were placed into the furnace where the temperature was first increased to 650 °C at a rate of 20°C/min and hold for 5 minutes. From 650 °C the temperature was then increase by a rate of 10 K/min to the brazing temperature 1100 °C where the dwell time was 15 min. After the dwell time, free cooling was initiated until room temperature.

Subsequently, the cemented carbide-steel joint with the carbon-hardening coldwork steel 1.2714 part was heated to a temperature of 850°C for 10 min by torch and then quenched to room temperature in oil. After that, tension relaxing was conducted in a vacuum furnace at 200°C for 2 hours.

Subsequently, the cemented carbide-steel joint with the carbon-hardening coldwork steel 1.2714 part was aged at 500°C for 2 h.

The sample will herein after be called Comparative 1.

Example 7

The samples were analyzed with respect to shear strength, surface hardness, core hardness and hardness depth curve.

The shear strength was analyzed by a shear device set-up as shown in Figure 3 where 1 is the steel part in the shape of a steel cylinder (0=20 mm, h=5 mm) and 2 is the cemented carbide part in the shape of a cemented carbide cylinder (0=10 mm, h=5 mm. The steel cylinder is positioned in the gap of the shear strength test device and therefore can only be moved in the loading direction. A notch, which was eroded into the surface of the device, held the joined parts in the right position and guarantees an evenly distributed force induction into the braze joint. The applied force, F, was constantly increased until the braze joint failed and the cemented carbide cylinder sheared off. The ultimate shear strength was then calculated by the quotient of the maximal measured force and the initial joining surface (A=78,5 mm ² ). The braze material was not removed before the determination of brazed joint shear strength.

To determine the depth of the nitriding according to Example 4, the average nitriding hardness depth was determined at room temperature. This was done by making a hardness depth curve by, on the transverse section of a nitrided sample, measuring HV 0.3 according to the standard DIN EN ISO 6507-1, starting with the first indent 0.025-0.1 mm from the edge, and then every 0.03-0.10 mm until the hardness did not change anymore. The hardness values obtained are recorded as a function of the distance from the surface. From this hardness curve, the nitriding hardness depth was taken as the distance between the surface and the limiting hardness (where the limiting hardness is the average core hardness (in HV 0.3) + 50HV 0.3). For Example 5, the nitriding depth was determined in the same way as Example 3, but with the difference that HV 0.5 was used. The core hardness given in Table 1, is in HV1 and has been measured by a Vickers hardness tester on a cross section of the maraging steel part, applying a load of 1 kgf (kilogram force) and a loading time of 15 s.

A pattern of 5 indents placed 1.5 mm apart was performed according to the standard and the value given in Table 1 is an average of the 5 indents. The surface hardness measurements were performed on the nitrided surface and where at least 5 indents placed 1.5 mm apart was performed and the value given in Table 1 is an average of the 5 indents. The measurements were performed using a Vickers hardness tester, applying a load of 1 kgf (kilogram force) and a loading time of 15 s.

Table 1

As can be seen in Table 1, the nitriding will create a surface with considerably higher hardness than the core which will lead to an increased wear resistance.