STUDY OF THE HARDNESS OF GRAY CAST IRON AS APPLIED TO THE CUTTING PROCESS

Yu. A. ROZENBERG and A. K. BAIKALOV

(Presented by Professor, Doctor A. M. Rozenberg)

An extensive series of experiments [9, 10] showed that the dynamics of the process of cutting gray cast iron in turning and milling has much in common with the machining of plastic materials, steels. The cutting forces, chip shrinkage, coefficient of friction, and built-up-edge geometry vary as functions of the cutting-condition parameters according to the typical curves obtained in machining steels. At the same time, cast iron, being a brittle material, is hardly work-hardened and does not produce a continuous chip.

To determine the magnitude of the stresses in the shear plane and to carry out a theoretical analysis of the chip-formation process, it is difficult here to use such important characteristics as chip shrinkage and its hardness. An indirect determination of chip shrinkage from the shear angle and calculation of the magnitude of the stresses in the shear plane confirmed the small degree of strengthening of cast iron. In this connection, the initial hardness of the brittle metal may be taken as its principal physicomechanical characteristic. On this basis, a theoretical-experimental formula was given for calculating the cutting force in machining cast iron [9, 10].

However, a number of questions remained unresolved. Since cast iron was not work-hardened, it was unclear what the role of the built-up edge was and why it was capable of cutting and of retaining a certain geometry. It was necessary to clarify more precisely the capacity of the metal for work hardening, and to explain why the hardness of cast iron can characterize the magnitude of the stress in the shear plane. In addition, cast iron, having in principle the same structural constituents as steel, possesses, in comparison with steel, a strong abrasive capacity. Experiments established that the tool lives of tools made of VK8 alloy, used in practice ($T = 60$—$240$ min), lie in a narrow temperature interval (400—600°), considerably lower than for steels (700—800°). Nevertheless, the cutting speeds for cast iron are 3—4 times lower than the speeds in cutting steels.

To clarify these questions, a study was carried out of the structure of cast iron on specimens of different hardness; hardness was measured by various methods at different ratios—

under loads relative to the ball diameter, the microhardness of the structural constituents, the hardness of the built-up edge and of the chip elements were determined.

Structure of the Cast Iron Studied

The structural constituents of the cast iron were studied on 21 polished sections cut from various cast-iron blanks in the hardness range from 140 to 300 Brinell units. After finishing, the polished sections were etched in a 5-percent solution of nitric acid in alcohol, examined, and photographed on a small MIM-6 metallographic microscope. Using the photographs, and directly on the frosted glass of the instrument, the percentage composition of the structural constituents (ferritic and ledeburitic regions) was determined by tracing the outlines of the grains on tracing paper until a figure regular in area was filled (mainly a quadrilateral). For a reliable determination of the percentage composition, the grains were traced over a large area. Thus, at an optical magnification of \(40^{x} \times 7.5^{x}\), the grain outlines were collected from 100 “passes” of the microscope field of view.

Study of the polished sections showed that the specimens contained 3 structures.

1. A structure with a ferrite–pearlite matrix, with a complete absence of ledeburitic grains. Ferrite content from 5 to 20% (Fig. 1).

2. A structure with a pearlitic matrix, with a complete absence of ferrite grains, but with a small proportion of ledeburitic grains (less than 0.2–0.5%) (Fig. 2).

3. A structure with a pearlite–ledeburite matrix. No ferrite; ledeburite from 0.5 to 4% (Fig. 3).

The pearlite of the structure is also nonuniform. It differs in grain size and in their shape. In most structures it is represented by coarse lamellar pearlite, alongside which there were structures with fine, swirling pearlite plates. In some specimens, within the general mass of the pearlitic matrix, isolated spots showed

Fig. 1

areas with very fine pearlite plates, which became noticeable only at magnifications greater than 500×. As a rule, these areas were distinguished by noticeably increased hardness.

Fig. 2.

Ledeburitic grains had the most varied shape and size: from small individual inclusions of more or less regular form to a continuous network of complex contour, part of which is shown in Fig. 3.

Fig. 3.

Most structures (18 out of 21) have ledeburitic inclusions, i.e., they belong to so-called mottled cast iron, obtained as a result of crystallization of the metal according to a mixed decomposition scheme [14]. The picture of graphite inclusions is very diverse. The size of the graphite inclusions varied within the range from 2 to 5 points.

The graphite inclusions themselves are nonuniform in size and shape. Three types of graphite inclusions can be distinguished: graphite inclusions in the form of macrocracks (with slit dimensions: width 2–10 µm, length up to tenths of a mm), inclusions in the form of microcracks located between the macrocracks (slit dimensions: width fractions of a micron, length up to ten microns), and separate globular inclusions from fractions of a micron to tens of microns. Usually microcracks are not visible on an unetched polished section and appear after light etching of the surface layer. Macro- and microcracks disrupt the integrity of the matrix and create an almost continuous network of voids, “loosening” the structure of the cast iron as a whole.

As a result of etching, the ledeburitic grains stand out most clearly in the structure (as white, irregularly shaped areas with sharp boundaries), as do ferritic grains, whose boundaries are also more or less clearly outlined and can be distinguished from the general field of the structure. The pearlitic constituent is also seen in the form of separate lamellae, but here it is impossible to distinguish the boundaries of an individual grain. Consequently, the pearlitic matrix constitutes the main “gray” background of the structure (by area from 70 to 100%), against which the remaining constituents are easily distinguished.

The chemical composition of the cast iron under study was determined only for some of the specimens and lay within the following limits:

\[ C_{\text{bond}} = 0.41\text{–}0.99\%;\quad \mathrm{Si}=1.15\text{–}2.65\%;\quad \mathrm{Mn}=0.25\text{–}0.76\%. \]

The other constituents were not determined.

Microhardness of the Cast-Iron Structure

The hardness of individual constituents of the cast-iron structure was determined on a PMT-3 hardness tester of the M. M. Khrushchev and E. S. Berkovich system, at loads \(P = 20, 50, 100\) grams. When the pyramid is pressed into a metallic matrix “loosened” by macro- and microcracks of graphite formations, a certain amount of graphite inclusions enters the volume of the indentation, which naturally distorts the measurement result toward a lower hardness compared with the continuous mass of the structure. A single measurement, and even a series of measurements, therefore cannot provide a reliable result. Several dozen indentations must be made in order to obtain a stable result. Because of the very low hardness of crystallization graphite \((H_v = 5\ \text{kg}/\text{mm}^2)\), when measuring microhardness, indentations in the graphite zone or in its immediate vicinity were not taken into account. With this measurement procedure it was easy to exclude from the measurement microcracks clearly visible on the polished surface of the specimen. However, microcracks and micropores located below the surface of measurement inevitably fell within the zone of penetration of the diamond pyramid and affected the measurement result. It is obvious that, in this case, the depth of penetration of the pyramid will affect the amount of graphite formations entering the volume of the indentation.

M. M. Khrushchev and E. S. Berkovich [12] recommend taking the size of the indentation to be considerably smaller than the dimensions of the grain being measured,

so that from the impression to the grain boundary there were no fewer than three diagonals. We followed these recommendations when measuring the hardness of ferritic and ledeburitic grains, whose boundaries are clearly outlined in the structure. As for the pearlitic matrix, the procedure for measuring its hardness was somewhat different and was carried out, for comparison, at three loads: \(P = 20,\ 50,\ 100\) grams. Impressions, numbering 50–150, were made with a uniform spacing of \(0.02\)–\(0.05\) mm, either over a continuous area, excluding only graphite inclusions, or in the form of a traverse over a length of 1–2 mm, including graphite formations. In the latter case, impressions in the graphite zone were not taken into account when calculating the averaged hardness, and their number served only as a measure for the quantitative accounting of graphite inclusions.

The overall average hardness of the structure as a whole was calculated by the formula:

\[ H_{\mathrm{av}}=\frac{S_{n}\cdot H_{n}+S_{\phi}\cdot H_{\phi}+S_{\ell}\cdot H_{\ell}}{100},\quad [\mathrm{kg/mm^2}]. \tag{1} \]

Here \(S_{n}; S_{\phi}; S_{\ell}\) are the areas, in percent, of pearlite, ferrite, and ledeburite; \(H_{n}; H_{\phi}; H_{\ell}\) are the average hardnesses of the corresponding structural constituents.

Formula (1) averages the microhardness of the specimen being measured and gives the same results as determining macrohardness with indentation by an indenter having a large impression area. The difference is that, when measuring average microhardness, graphite inclusions were excluded from the measurement. Thus, the averaged microhardness characterizes only the hardness of the metallic matrix of the structure, whereas macrohardness characterizes the structure as a whole.

To clarify the influence of the nonmetallic part of the structure on the hardness characteristic, a comparison was made of the results of hardness measurements at different impression areas, which included graphite inclusions to different degrees.

For this purpose, a Vickers hardness tester was used (diamond pyramid at \(P=5\) kg, steel ball 2.5 mm in diameter at \(P=62.5\) kg) and a Brinell press (\(P=3000\) kg, steel ball \(\varnothing=10\) mm). When measuring hardness, the impression diagonal or impression diameter varied over very wide limits (diagonal from \(0.008\) to \(0.7\) mm, impression diameter from \(0.6\) to \(4.5\) mm). Comparative data from measurements of the microhardness of the pearlitic mass and the microhardness of the entire structure are presented in Fig. 4.

The graph confirms the great influence of graphite formations on the result of hardness measurement, which is expressed in the substantial difference between the hardness of the metallic matrix (microhardness) and the hardness of the entire structure (macrohardness), and also in the tendency for hardness to decrease as the area (depth) of indenter penetration increases, since in this case a relatively larger number of graphite inclusions enters the impression zone.

Thus, in studying questions of tool wear and life, which will be determined mainly by the abrasive capacity of the metallic matrix, one should proceed from the hardness determined by microimpressions with the smallest indenter penetration depth. Obviously, impressions with a load of \(P=20\) grams are, from this point of view, the most acceptable. The macrohardness of cast iron, which depends strongly on the size and shape of graphite inclusions, cannot be a reliable characteristic of the abrasive capacity of the metallic matrix of cast iron.

The difference in the hardness of the metallic matrix and of the structure as a whole is very substantial (a change in hardness by a factor of 2–2.5), and the greater, the more graphite inclusions the structure contains. Thus, for specimen No. 7, characterized by very large graphite inclusions (score 2), the hardness varies from 187 to 490 units, i.e., by a factor of 2.6, whereas for specimen No. 4, which has fine graphite inclusions (score 5), the hardness difference is 180–310 units, i.e., only by a factor of 1.7.

Such a substantial difference in the hardnesses of the metallic matrix and of the structure as a whole explains the increased abrasive capacity of cast iron observed in experiments, as compared with steels.

Fig. 4.

With the same macrohardness of cast iron and steel, the hardness of the metallic matrix of cast iron will be 2–2.5 times higher, which in cutting at speeds corresponding to the same tool life will inevitably cause a substantial lowering of the cutting speed when machining cast iron. Naturally, the cutting temperatures in cast irons will in this case also be somewhat lower than in steels. In essence, cutting cast iron of the same macrohardness as steel is equivalent to cutting hardened steels. From this point of view, it would be expedient to compare the machinability of cast irons and steels having the same hardness of the metallic matrix.

To clarify the abrasive capacity of the metallic matrix of cast iron, an experiment was carried out on turning certain cast irons with different hardness and structure.

For the purpose of carrying out a tool-life experiment within the limits of one cast-iron bushing, characterized by a uniform structure, the following cutting conditions were selected: \(V = 100\ \text{m/min}\), \(t = 2\ \text{mm}\), \(S = 0.23\ \text{mm/rev}\). The blunting criterion was taken as equal to \(1\ \text{mm}\). Cutters made of VK8 alloy were sharpened with the following geometry:

\[ \gamma = 10^\circ,\quad \alpha = 10^\circ,\quad \alpha_1 = 8^\circ,\quad \varphi = 45^\circ,\quad \varphi_1 = 15^\circ,\quad r = 0.5\ \text{mm}. \]

The results of the experiment for determining the averaged hardness of the metallic matrix and its influence on the life of carbide cutters are given in the table.

The results of the experiment are presented in Fig. 5. Here the life-

Table 1

tool life is related to the hardness of the metallic matrix (upper curve) and to the hardness of the structure as a whole (lower curve). It follows from Fig. 5 that the microhardness of the metallic matrix, for most structures, is more or less regularly correlated with tool life and, apparently, characterizes the wear-resisting capacity of cast iron. The macrohardness of cast iron does not show such regularity. The experimental points for bushings Nos. 7, 8, and 12, which have large graphite inclusions, “failed” in terms of hardness. However, for the upper curve, the machinability of the structure of specimen No. 13 clearly falls outside the general regularity. Its structure is characterized by large differences from the other specimens (Fig. 3).

Fig. 5.

For analysis of the nature of the change in hardness of the metallic matrix of this specimen, Fig. 6 presents its traverse hardness, which differs substantially from specimens 8, 4, and 12, whose traverse hardness is shown in Figs. 7 and 8. Here the ordinate gives the hardnesses of individual impressions, made at a spacing of 0.02 mm, in a number of 50–100. Penetration of the indenter into graphite inclusions is represented in the figures by zero hardness.

On the basis of Fig. 3, Fig. 6, and the table, it may be concluded that the structure of specimen No. 13 is characterized by a higher content of ledeburitic regions of high hardness \((H_l = 1200–1350\ \text{kg/mm}^2)\) and by a very great nonuniformity of hardness of the metallic matrix. Thus, the hardness of the pearlitic matrix lies within the range

Fig. 6.
Within the figure: Sample No. 13.
Ordinate: Hardness $H_n$, kg/mm$^2$.
Scale marks: 200, 300, 400, 500, 600, 700, 800, 900, 1000, 1100, 1200.
Horizontal length: 2 mm.

Fig. 7.
Within the figure: Sample No. 8.
Ordinate: Hardness $H_n$, kg/mm$^2$.
Scale marks: 200, 300, 400, 500, 600, 700, 800.
Horizontal length: 2 mm.

Fig. 8.
Within the figure: Sample No. 4; Sample No. 12.
Ordinate: Hardness $H_n$, kg/mm$^2$.
Scale marks: 200, 300, 400, 500, 600, 700, 800.
Horizontal lengths: 1 mm; 1.28 mm.

…300–400 units; however, against this background there are more or less uniformly distributed regions of increased hardness \((H_p = 500\text{–}580\ \mathrm{kg/mm^2})\), which are clearly visible under the microscope as spots with finely swirled and poorly resolved pearlite on the plate. In addition, near the ledeburitic grain there is a zone, approximately equal to the area of the grain, with still higher hardness \((H_p = 700\text{–}800\ \mathrm{kg/mm^2})\). Taking all these deviations of hardness into account and summing them by formula 1 gives an average hardness \(H_{\mathrm{av}} = 454\ \mathrm{kg/mm^2}\). This hardness is clearly insufficient to explain the increased wearing capacity of the structure.

It is obvious that structures with a high percentage of ledeburite content (25% and more) and with the presence of regions of high hardness embedded in a “soft” pearlitic matrix cannot, from the standpoint of wearing capacity, be characterized by the average hardness. For pearlitic cast irons with a low ledeburite content (1% and less) and with a more stable hardness of the metallic matrix, the average microhardness may be accepted as a measure of machinability.

For final conclusions, the accumulation of a larger amount of experimental data is required.

Dependence of Cutting Forces on the Hardness of Cast Iron

The Brinell hardness value is introduced into calculation formulas for determining the cutting force when machining gray cast iron. Usually these formulas have the empirical form:

\[ P_z = C \cdot H_{\mathrm{B}}^n, \tag{3} \]

where the values of \(n\) are determined from experimental data and differ among different investigators. Thus, according to Kronenberg [2], \(n = 0.4\); according to Bezprozvannyi [1], \(n = 0.53\); according to Danielyan [4], \(n = 0.55\); according to the 1950 standards [6], \(n = 0.55\); according to the 1958 standards [8], \(n = 0.4\). The reason for the discrepancy in these data is the incorrect method, used by the above-mentioned investigators, for determining the relationship between forces and hardness. All the investigators determined the dependence of the cutting force on the hardness of the cast iron being machined at a constant cutting speed and thereby did not take into account the influence of cutting temperature on the cutting forces.

With an increase in the hardness of cast iron, the cutting temperature rises (provided the other factors of the cutting regime remain constant), as a result of which the cutting forces are affected not only by the factor of the change in hardness, but also by the factor of the change in temperature. Therefore, the only correct method for determining the dependence of cutting forces on Brinell hardness is the method of determining this dependence at constant temperature.

In this case it is found [10] that the cutting forces are directly proportional to the Brinell hardness of cast iron, i.e., the coefficient \(n = 1\), and

\[ P_z = C \cdot H_{\mathrm{B}}. \tag{4} \]

This relation already has a certain physical meaning, since, according to a number of investigators [9], measurement of hardness by indentation reveals not some special mechanical property of the metal, but the same resistance to plastic deformations that is manifested in other types of plastic deformation,—

tension, compression, torsion, etc. This same property should also manifest itself in cutting, since the cutting process is a deformation of simple shear.

On the basis of this, and of N. N. Davidenkov’s proposition [3] that the concept of indentation hardness is based on stress, L. A. Khvorostukhin [11] experimentally established a single relationship between hardness and stress in cutting for various machined materials (steel, copper, aluminum), and also showed that this relationship is valid for other schemes of plastic deformation (compression, tension, torsion):

\[ \tau = 0.185\,H_n, \tag{5} \]

where, for the cutting process:
\(\tau\) is the shear stress in the shear plane,
\(H_n\) is the hardness of the removed chip, determined with a TG device using a diamond pyramid, under a load of \(5\text{–}10\) kg.

An analogous relationship between stress and hardness was obtained [10] for gray cast iron both in cutting and in compression:

\[ \tau = 0.18\,H_B, \tag{6} \]

where \(H_B\) is the Brinell hardness of the original material.

At the same time it is known that the numerical values of ball hardness for a given specimen depend on the magnitudes of the load and the diameter of the ball [7]. Therefore, for a physical substantiation of the above relationship between \(\tau\) and \(H_B\), it is necessary to verify the correctness of determining the hardness of gray cast iron at \(P = 3000\) kg and \(D = 10\) mm.

In order to obtain the same Brinell hardness number with balls of different sizes, it is necessary in testing to use loads proportional to the squares of the ball diameters. In the present work the dependence of the hardness \(H_B\) of gray cast iron on the ratio \(\frac{P}{D^2}\) was determined. Different values of this ratio were obtained by varying the load from 187.5 to 13,500 kg at a constant ball diameter of 10 mm. For comparison, not only gray cast iron but also steel 45 and malleable cast iron were subjected to testing.

The test results are given in Fig. 9. From this figure it is evident that the numerical value of the hardness of gray cast iron does indeed depend strongly on the ratio \(\frac{P}{D^2}\). With an increase in the ratio \(\frac{P}{D^2}\), the hardness at first rises sharply, then over a wide range of changes in \(\frac{P}{D^2}\) remains constant, and then decreases.

The decrease in hardness at high values of \(\frac{P}{D^2}\) was accompanied by the appearance of small cracks around the impressions. At \(\frac{P}{D^2} = 25\text{–}105\) the numerical value of the hardness is constant, maximal, and, according to Meyer, corresponds to the limiting ball hardness number, which, in accordance with [7], does not depend on the diameter of the ball used. Conse-

Consequently, using in the tests

\[ \frac{P}{D^2}=30, \]

we obtained the limiting hardness value.

As can be seen from Fig. 9, the limiting hardness value for malleable cast iron was obtained at

\[ \frac{P}{D^2}=35\text{--}40, \]

and for steel 45 no limit was obtained at all.

N. N. Davidenkov [3] put forward a proposal for the indirect determination of various mechanical properties from hardness tests. In recent years this proposal has been widely developed in the works of Soviet scientists. Thus, M. P. Markovets [5] established a relationship between hardness under ball indentation and the tensile strength for various metals, including cast iron. He showed that, in determining this relationship, it is necessary that the degrees of deformation in determining hardness and in tension coincide. In this case

Fig. 9.
Vertical axis: Brinell hardness. Horizontal axis: ratio \(\frac{P}{D^2}\).
Legend: ○ — steel 45; × — gray cast iron; □ — malleable cast iron.

\[ \sigma_{\mathrm{в}}=0.33H_{\mathrm{Б}}. \tag{7} \]

Since for gray cast iron the degree of deformation in tension is considerably lower than in hardness measurement, this relationship for it has a different form:

\[ \sigma_{\mathrm{в}}=0.15H_{\mathrm{Б}}. \tag{8} \]

This resulted from the fact that for such an inhomogeneous material as gray cast iron it is impossible to determine the relationship between Brinell hardness and tensile strength. The Brinell hardness of cast iron is a macrohardness that includes the hardnesses of all the constituents of the cast iron, including the hardness of the graphite inclusions. In a tensile test of cast iron, its tensile strength is determined by the strength not of all its constituents. In particular, the strength of the graphite inclusions drops out, and therefore

the attainable degrees of deformation in tension of gray cast iron are very small.

In connection with this, in our opinion, it is necessary to seek a relation between hardness and the ultimate strength in compression, and not in tension. In determining hardness and in compressing cast iron, the attainable degrees of deformation are approximately equal, and cast iron, as a nonuniform material, evidently behaves in the same way in these cases.

Markovets studied in detail the deformations obtained on the surface of the indentation when a ball is pressed in. He showed that the degree of deformation on the surface of the indentation is distributed nonuniformly, and an experimental curve was constructed showing the relation of the average degree of deformation in the indentation \(\psi_{cp}\) to the ratio of the indentation diameter \(d\) to the ball diameter \(D\). On the basis of experiments and the data of this relation, we determined the dependence of \(\psi_{cp}\) on the ratio

\[ \frac{P}{D^2} \]

for various metals.

Using this dependence and the experiments carried out on compression of cast iron, we constructed a graph (Fig. 10) relating the initial stresses to the average deformation in measuring hardness and in compression.

Fig. 10.

compression. The tangential stresses were determined in compression as \(\tau = 0.5\sigma\), and in hardness measurement as \(\tau = 0.18H_B\).

As can be seen from Fig. 10, the values of the tangential stresses determined in hardness measurement and in compression of cast iron, at equal degrees of deformation, agree well, especially at \(\psi_{cp} \ge 8\%\). At \(\psi_{cp} \ge 8\%\), the normal stress \(\sigma\) is already the ultimate strength of cast iron in compression, and the relation between \(\sigma_b\) and \(H_B\) is determined by the formula

\[ \sigma_b = 0.36 H_B . \tag{9} \]

In its numerical coefficient this relation is analogous to the relation established by Markovets for steels and other materials. Since in cutting the degrees of deformation are greater than 8% and the cutting process is identical to the compression process, i.e., in this case the properties of cast iron are determined by the limiting (maximum attainable) value of its macrohardness, the obtained dependence between stress and hardness (6) is ...

is not accidental. Consequently, the stresses during cutting and the cutting forces must be directly proportional to the Brinell hardness of the cast iron, determined at

\[ \frac{P}{D^{2}} > 25 . \]

In deriving relation (6) for cast iron, its initial hardness was taken, not the hardness of the chip, since it was assumed that cast iron is a metal not subject to work hardening. Obviously, the capacity of cast iron for work hardening had to be checked. The work-hardening ability of gray cast iron was investigated by the method of hardness measurement. A ball with \(D = 10\) mm was pressed in under various loads. In this case different ratios \(\frac{P}{D^{2}}\) and \(\frac{d}{D}\) were obtained and, consequently, different degrees of deformation in the indentation. Then the hardness of the surfaces of these indentations was measured on a Vickers instrument with a ball \(D = 2.5\) mm under a load \(P = 62.5\) kg. Under the same conditions the initial hardness of the material was also determined. Comparison of the initial hardness with the hardness of the indentation surfaces made it possible to determine the degree of increase in hardness, i.e., the degree of work hardening. Figure 11 shows the dependences of work hardening in the indentation on the average degree of deformation in the indentation for various metals. As is seen from this figure, the maximum work hardening of gray cast iron is 10–12% and is reached at an average degree of deformation \(\psi_{\mathrm{av}} = 10\%\); with a further increase in \(\psi_{\mathrm{av}}\) it remains unchanged. Consequently, under conditions of hardness measurement close to conditions of triaxial compression, the maximum work hardening of cast iron is no more than 10–12%. During cutting we have, in the main, a plane deformation scheme, far removed from the conditions of triaxial compression, and, consequently, the work hardening of cast iron in this case will be still less and may be neglected.

Fig. 11.

It is also seen from Fig. 11 that the maximum work hardening of malleable cast iron is 38%, while for steels 45 and EIIT the work-hardening limit has not been established, which corresponds to reality and substantiates the correctness of the method we selected for determining work hardening.

Hardness of built-up edge, build-up, and chip and their comparison with the initial hardness of cast iron

Determining the hardness of the built-up edge is of great importance for studying built-up-edge formation and for determining the role of the built-up edge in the cutting process. Measurement of the hardness of the built-up edge when cutting steels is performed on polished sections obtained with the aid of a “falling tool” device. Such a method, when cutting cast iron, cannot be applied because of its brittleness. Therefore, measurement of the hardness of built-up edges,

in cutting cast iron in the present work was carried out directly on the tool. After being removed from under the chip, the carbide insert together with the built-up edge was taken out of the holder and placed on the stage of a PMT-3 hardness tester. Because of the small dimensions of the built-up edge, it was possible to measure only its microhardness. For reliability of the measurements, 10–12 impressions were made on each built-up edge under loads of 20, 50, and 100 g. Figure 12 shows the dependence of the microhardness of the built-up edge on the cutting temperature on the rake face of the tool. The same figure shows the microhardness of the accretion, the chip, and the original material. To determine the microhardness of the chip, its elements were pressed into wooden blocks and subjected to careful grinding and polishing. The microhardness of the original cast-iron material was determined by the traverse method (25 impressions each under loads of 50 and 100 g). During measurement, graphite formations were excluded.

As can be seen from Fig. 12, the microhardness of the built-up edge and of the accretion changes with increasing cutting temperature according to a definite law.

With increasing temperature, the microhardness of the built-up edge rises sharply and reaches a maximum at a temperature of 300°. The maximum hardness of the built-up edge corresponds to the maximum value of the front angle of the built-up edge.

Fig. 12.

At the same time, the microhardness of the built-up edge reaches a very large value, \(H_n = 900\text{–}1000\). With a further increase in temperature, the hardness of the built-up edge decreases and reaches a minimum at a temperature of 550°, at which the built-up edge disappears. The microhardness of the accretion increases with increasing temperature and at temperatures of 800–1000° reaches values of \(H_n = 800\text{–}1000\). At these temperatures, on the flank face of the tool a built-up edge is formed in the form of long locks with a hardness equal to the hardness of the accretion on the rake face, i.e., \(H_n = 800\text{–}1000\).

Measurement of the microhardness of the chip showed that its microhardness does not depend on the cutting conditions and exceeds the initial microhardness of the cast iron by 5–8% (Fig. 12). At the same time, as was shown earlier, the maximum value of the hardness of the built-up edge and of the accretion is 2.5–3 times higher than the initial microhardness of the cast iron. Such high hardnesses can evidently be explained either by the fact that, in the structure of the built-up edge and accretion, the principal constituent is the hardest constituent of cast iron—ledeburitic inclusions, having a hardness of \(H_n = 800\text{–}1400\)—or by the conditions of formation of the built-up edge and accretion (high pressures and temperatures), under which large strengthening of individual constituents of the cast iron (ferrite and pearlite) or even structural transformations are possible. The hardness of the chip, exceeding the initial hardness of the cast iron by 5–8%, is perhaps nevertheless the result of the presence of strengthening in the cutting process. Evidently,

In order to give a complete explanation of these phenomena, it is necessary to conduct studies of the possibility of strengthening the individual constituents of gray cast iron.

Conclusions

Gray cast iron, by its composition, is a heterogeneous material that includes a metallic matrix and nonmetallic graphite inclusions. In this connection, its mechanical characteristics should include both the macrohardness and the microhardness of its individual constituents. In determining the relationship between these characteristics and the phenomena occurring during the cutting of cast iron, it is necessary to distinguish between the hardness of the metallic matrix of the cast iron and the overall (total) hardness of the structure as a whole, since in the process of tool wear the abrading capacity of cast iron is determined by the structural constituents of the metallic matrix, whereas in chip formation the strength characteristics of the structure as a whole come to the fore.

On the basis of these premises, the authors obtained relationships between tool life and the microhardness of cast iron, between cutting forces and the macrohardness of cast iron, determined the possible strengthening of cast iron in the macrovolume, and determined the hardness of the built-up edge, deposit, and chip in comparison with the initial hardness of the cast iron.

References

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