SURFACE FINISH IN TURNING OF GRAY CAST IRON

Yu. A. Rozenberg, V. M. Baganov, O. A. Kudryavtsev

(Presented by Prof. A. M. Rozenberg, Doctor of Technical Sciences)

The problem of the quality of the finish of a machined surface is one of the most important problems of modern mechanical engineering. However, this problem still contains a number of insufficiently resolved questions. This applies especially to surface finish in the machining of gray cast iron. The information available in the literature on surface finish in the machining of gray cast iron is fragmentary and does not quite accurately outline the true picture. Thus, in the works of Dolmatovsky [2], O. G. Beletsky [3], P. E. Dyachenko, and M. O. Jakobson [4], it is stated that in machining gray cast iron by turning and face milling methods, a finish of ∇∇∇ 7—∇∇∇ 8 is quite attainable. In the departmental standards of the Ministry of the Coal Industry, VN—137, in accordance with GOST—2789—51, a finish of ∇∇ 6 is required on explosion-proof surfaces of fixed joints of parts of electric mining machines.

However, at the Tomsk Electromechanical Plant named after V. V. Vakhrushev, under shop conditions, it was not possible in face milling to ensure reliable attainment of surface finish ∇∇ 6. Engineer A. V. Pozolotin of this plant carried out a special small investigation, which did not confirm the data from the above-mentioned works. In connection with this it was necessary to decide what class of finish is economically attainable in the machining of gray cast iron. Most investigations of surface finish have been carried out in cutting steels. These works made it possible to establish a strict regularity in the change of surface finish in the machining of steels. Analyzing the works of many investigators, and also basing himself on his own experiments, A. N. Eremin [5] comes to the conclusion that in non-free cutting of steels the surface finish is characterized by irregularities whose magnitudes are composed of two elements according to the equation

\[ H = H_{theor} + H_{sher}, \]

where \(H_{theor}\) is the geometrical characteristic of surface finish, depending on the profile of the tool in plan \((\varphi, \varphi_1 \text{ and } r)\), the condition of the cutting edge, and the magnitude of the feed.

\(H_{sher}\) is a geometrical and at the same time physical characteristic of the finish of the cut, depending on the intensity of the temperature regime during cutting and on the plastic properties of the material being machined.

Further, A. N. Eremin indicates that, as with all phenomena of the cutting process (built-up edge, shrinkage and hardness of chips, coefficient of friction, etc.), surface finish changes under the influence of the cutting temperature as a result of a change in \(H_{sher}\). The cutting temperature affects the magnitude of

$H_{\text{rough}}$ through changes in friction on the flank face of the cutter or in the built-up edge, the degree of plastic deformation, i.e., chip shrinkage, the height of the built-up edge on the cutter, and shears on the cut surface under the cutter.

The general regularity of the change in surface quality when cutting gray cast iron, and the physical substantiation of this regularity, have not been revealed, owing to the small number of investigations on this question. In these investigations one can find only studies of individual dependencies of $H$ on the parameters of the cutting process and certain assumptions concerning their physical basis. Thus, for example, P. E. Dyachenko, in his work [6], indicates that in cutting gray cast iron the curve of the dependence $H_{\max}=f(v)$ has a hump-shaped character. From this he makes the assumption that in cutting gray cast iron “a small built-up edge is nevertheless obtained.”

In the Metal Cutting Laboratory of the Tomsk Polytechnic Institute named after S. M. Kirov, Yu. A. Rozenberg carried out work on investigating the process of cutting gray cast iron [7]. In this work a number of regularities were revealed and a physical substantiation of the dynamics of the process was given. In particular, it was found that when cutting gray cast iron a built-up edge is formed and that all the phenomena of the cutting process of gray cast iron (the built-up edge, which changes the actual geometry of the tool, the coefficients of friction on the rake and flank faces of the tool, shrinkage, etc.), just as in the cutting of steels, change under the influence of the cutting temperature (Fig. 1). Apparently, these revealed regularities can make it possible to give a more complete substantiation of the dependence of surface finish in machining gray cast iron on the parameters of the cutting process.

Fig. 1

All that has been set forth above prompted the authors to conduct an investigation of surface finish in the turning of gray cast iron. The work was carried out under an agreement on creative collaboration with the Tomsk Electromechanical Plant named after V. V. Vakhrushev in the Metal Cutting Laboratory of the Tomsk Polytechnic Institute named after S. M. Kirov.

Research Method

The theoretical height of the residual ridges of the ideal surface profile $H_{\text{theor}}$ is an exclusively geometrical quantity, depending on the feed $S$, the nose radius of the cutter $r$, and the plan angles $\varphi$ and $\varphi_1$. In the literature one can find a number of equations that mathematically express the indicated dependence. Thus A. I. Isaev [8] gives the following equations, which express the value $H_{\text{theor.max}}$ for four main cases of obtaining the micro-roughness profile when measuring it in the feed direction.

In finishing operations, for which $S<0.2$ mm/rev and $r>0.5$ mm, the value $H_{\text{theor.max}}$, as is evident from the table, is determined by the formula:

IV case, when the entire profile of the microroughnesses on both sides is formed by an arc of radius $r$.

Table 1

In this case, a change in the plan angles $\varphi$ and $\varphi_1$ has no influence on the microroughness profile from the geometrical standpoint. A change in the angles $\varphi$ and $\varphi_1$ affects the microroughness profile through changes in cutting temperature, tool life, and frictional forces on the rear face of the tool.

In view of this, in our investigations the geometry of the tool was selected on the basis of its optimum life.

As the cutting tool, turning tools fitted with plates of hard alloy VK–8 were used, with the following geometrical parameters:

Main plan angle $\varphi = 45^\circ$.

Auxiliary plan angle $\varphi_1 = 13 \div 15^\circ$.

Rake angle $\gamma = +10 \div -12^\circ$.

Main clearance angle $\alpha = 8 \div 15^\circ$.

Auxiliary clearance angle $\alpha_1 = 10^\circ$.

Radius of rounding of the tool nose $r = 0 \div 2$ mm.

Before each experiment the faces of the cutting tool were sharpened and carefully lapped on a cast-iron disk with boron carbide powder. The angles of the tools were measured on a universal angle gage with an accuracy of up to $30'$. When setting the tool geometry, it was checked with an angle gage. The amount of wear and the radius of rounding of the tool nose were measured with a Brinell portable microscope.

All experiments were carried out on two lathes of medium rigidity. The following cutting conditions were used: the cutting speed varied within the range from 7 to 400 m/min, the feed $S$ varied within the range from 0.03 to 0.6 mm/rev, and the depth of cut varied within the range from 0.5 to 4 mm.

The range of cutting speeds used in the experiments covers all speeds encountered in practice. The choice of feeds was governed by the capabilities of the machines and by the task set before the authors: to find the economically attainable class of finish. All experiments were conducted with constrained cutting, without lubrication or cooling. The cutting speed was measur-

was carried out by measuring the number of spindle revolutions during the experiment with a tachometer. All experiments were conducted on modified gray cast iron with Brinell hardness \(H_B = 207 \div 217\). The cast-iron blanks were cast at the Tomsk Electromechanical Plant in the form of bushings of size \(l \times d_{\text{nar}} = 280 \times 118\) mm. Before the experiments, the scale was removed from the blanks. In deciding the question of the attainable class of surface finish, it was necessary to consider the surface finish when cutting with a worn tool. For this purpose the authors conducted two series of experiments to determine the influence of tool wear on surface finish in turning gray cast iron. Tool wear in finish cutting of cast iron is characterized mainly by abrasion of the rear faces of the tool. In their experiments the authors brought the wear of the tool on the rear face up to \(h = 1.1\) mm. Because of the shortage of the material being machined, the following experimental procedure was adopted. The tools were worn by the natural method to specified amounts of wear at a cutting speed \(V = 85\text{--}95\) m/min and feed \(S = 0.2\) mm/rev. Then, with the amount of wear obtained under this regime, the influence of wear on surface finish was determined under other cutting regimes. In this, the feeds were varied within the limits from 0.07 to 0.2 mm/rev and the cutting speed within the limits from 50 to 200 m/min. Thus, 12 experiments were carried out with each dulling.

In order to take into account the wear of the tool during each experiment, wear measurements were made after each experiment. From the results of the experiments, graphs were constructed of the dependence of \(H_{\max}\) on wear at different cutting speeds and feeds. These dependences in the first and second series of experiments were repeated well, with a difference in readings of 2–4 microns, i.e., within one and the same class.

The height of the irregularities \(H_{\max}\) in all experiments was measured with a Linnik double microscope, model MIS-III. \(H_{\max}\) was determined as the arithmetic mean of 10–12 measurements. The authors believe that the most accurate measurement of surface finish would have been obtained by using a double microscope and compiling profilograms. However, the absence of profilographs in the laboratory did not permit this to be carried out. In this connection, to verify the results obtained, the authors repeatedly repeated the experiments and obtained good agreement of results. In addition, the authors analyzed the scatter of the readings of 10–12 measurements in each experiment. As Fig. 2 shows, this scatter usually does not exceed 15–20%.

Fig. 2

In all, more than 400 experiments and more than 4000 measurements were carried out in the work.

Results of the investigations

Influence of cutting speed and feed on surface finish

As can be seen from Fig. 3, the dependence \(H_{\max}=f(v)\) at different feeds has the same character. With an increase in cutting speed, \(H_{\max}\) decreases. The curve of the dependence \(H_{\max}=f(v)\) has two sharply expressed zones of speed influence, within which the irregularities \(H_{\max}\) undergo different changes.

Fig. 3

In the first zone, an increase in cutting speed up to 40–60 m/min entails a sharp decrease in irregularities. In the second zone, from a cutting speed of \(v=40\text{–}60\) m/min, an increase in cutting speed leads to an insignificant decrease in irregularities. As can be seen, the curve of the dependence \(H_{\max}=f(v)\) in cutting gray cast iron differs from the same curve in cutting steels in that it does not have a sharply expressed hump. The character of the curve \(H_{\max}=f(v)\) in cutting gray cast iron can be explained on the basis of work [7]. As this work shows, in cutting gray cast iron the phenomenon of a built-up edge on the rake face of the tool is observed. The built-up edge, forming on the rake face of the tool, changes the geometry of the tool and thereby changes the cutting conditions. The geometry of the built-up edge depends only on the cutting temperature on the rake face of the tool and on the rake angle of the tool. The greater the feed, the greater the height of the built-up edge. With an increase in cutting speed (or, more correctly, cutting temperature), the rake angle of the built-up edge initially increases, reaches a maximum, and then decreases until the built-up edge disappears. The phenomenon of built-up-edge formation in cutting cast iron, as in cutting steels, has an adverse effect on the finish of the machined surface. Because the tip of the built-up edge, projecting far ahead of the tool cutting edge and below the cutting line, produces deep grooves and tears on the cut surface, the surface finish in the built-up-edge formation zone deteriorates. In addition, in the zone of low temperatures (that is, in the zone of low cutting speeds) the built-up edge is unstable. In this case the built-up edge, or part of it, is constantly torn off and passes along the flank face of the tool, as a result of which the coefficient of friction on the flank face of the tool increases and the quality of the machined surface deteriorates.

Thus, as can be seen from Fig. 4, the initial sharp decrease in the height of the irregularities with increasing cutting speed can be explained by a decrease in the coefficient of friction on the flank face of the tool and the disappearance of the built-up edge. Since the change in the geometry of the built-up edge and in the value of the coefficient of friction on the flank face is determined by the cutting temperature, it may be concluded that the inflection point of the curve \(H_{\max}=f(v)\), from a more intensive change in \(H_{\max}\) to a less intensive change, is also determined by the cutting temperature; that is, the smaller the feed, the higher the cutting speed corresponding to this point. This conclusion is confirmed by Fig. 3.

Fig. 4

According to the data of work [7], the temperature corresponding to the disappearance of the built-up edge, and consequently also the corresponding inflection point, is approximately \(500^\circ\). From all this it follows that the influence of feed, depth of cut, tool geometry, and other cutting-condition parameters on \(H_{\max}\) should be determined not at a constant cutting speed, but at a constant cutting temperature. Figure 5 shows the dependence \(H_{\max}=f(S)\) at a constant temperature \(T=500^\circ\). As can be seen from this figure, with decreasing feed \(H_{\max}\) decreases; moreover, at small feeds, beginning with \(S=0.05\) mm/rev, \(H_{\max}\) in fact remains unchanged. A similar pattern for the dependence \(H_{\max}=f(S)\) was obtained by S. P. Semenov [9]. The feed above which the zone of sharp increase of microroughnesses lies is called, for cast iron, the critical feed. In S. P. Semenov’s work the critical feed was equal to \(0.12\) mm/rev. It may be concluded that, at feeds below the critical value, the height of the irregularities is determined not by the feed, but by the radius of rounding of the tool tip, the rigidity of the machine, friction on the flank face of the tool, etc. Apparently, the magnitude of the critical feed itself also depends on these quantities.

Figure 5 includes, for comparison, the curve of the dependence \(H_{\text{theor}}=f(S)\) at \(r=0.5\) mm, where \(H_{\text{theor}}\) was determined by the formulas of A. I. Isaev. The comparison shows that at large feeds \(H_{\text{theor}}\) is greater than the actual height of the microroughnesses. Apparently, owing to the brittleness of gray cast iron, at large feeds the sharp protrusions of microroughnesses of geometric character are destroyed and smoothed out, as a result of which \(H_{\max}\) decreases. In addition, the comparison in Fig. 6 of the two curves confirms the preceding conclusion that, at small feeds, \(H_{\max}\) is determined mainly not by the feed, but by other quantities.

Fig. 5

Finally, analyzing the data from the experiments presented, one may conclude that in finishing operations it is necessary to work with cutting conditions corresponding to cutting temperatures above \(500^\circ\).

According to the data of work [7], cutting temperatures above \(500^\circ\) correspond to cutting conditions at a tool life \(T \leq 90\) min. In this case, when working with a sharp tool with \(r = 0.5\) mm at feeds \(S < 0.07\) mm/rev, the 6th class of finish can be obtained. When working with feeds in the range \(S = 0.07\text{–}0.3\) mm/rev, the economically attainable class of finish (under the given conditions) is \(\nabla\nabla 5\).

Influence of Depth of Cut on Surface Finish

Many investigators believe that the depth of cut has little influence, or no influence at all, on the magnitude of irregularities on the machined surface. Thus, for example, V. A. Rog [10] states that a depth of cut within the range \(0.1\text{–}2\) mm and more does not affect the microgeometry of the surface. The authors consider that such a view of the influence of depth of cut on surface finish is not entirely correct.

A change in depth of cut entails a change in cutting temperature. At the same time, as indicated above, the influence of depth of cut on the height of microirregularities should be determined at a constant temperature. Thus, if at a constant cutting speed the depth of cut is increased, the cutting temperature will rise, and consequently the surface quality should improve. Indeed, Fig. 6 shows that at a constant cutting speed, with increasing depth of cut, \(H_{\max}\) decreases. If, in Fig. 6, dependences \(H_{\max} = f(T^\circ)\) were plotted from the same experimental data, then the points in this figure for different depths of cut would approach one another. Thus, in the above conclusions on the influence of depth of cut on surface finish, a correction must be introduced—the condition of constant temperature.

Fig. 6

In addition, the authors consider that increasing the depth of cut within the limits permitted by the rigidity of the machine leads to greater stability of the cutting process, as a result of which the finish of the machined surface is somewhat improved.

Influence of Tool Geometry on Surface Finish

The principal parameters determining the geometry of the tool are the rake angle \(\gamma\), the clearance angles \(\alpha\) and \(\alpha_1\), the plan angles \(\varphi\) and \(\varphi_1\), and the nose radius of the cutter \(r\). A change in these parameters entails, to one degree or another, a change in the cutting temperature in the contact zone, which is reflected in the surface finish of the workpieces being machined. In addition, all these parameters of tool geometry may influence surface finish as geometrical characteristics. In assessing the influence of the parameters of tool geometry on surface finish as geometrical characteristics, the following conclusions may be drawn. A change in the rake angle of the tool is not reflected in its profile, but leads, as shown in [7], to a change in the height of the built-up edge. The greater the rake angle, the smaller the height of the built-up edge and, consequently, the smaller the height of the irregularities. Thus, by reducing the rake angle of the tool, we raise the cutting temperature and increase the height of the built-up edge. An increase in temperature leads to a reduction in the height of the irregularities, while an increase in the height of the built-up edge leads to an increase in the height of the irregularities. Indeed, owing to this dual action, as Fig. 7 shows, changing the rake angle from \(+8^\circ 30'\) to \(-12^\circ 30'\) in fact has no effect on surface finish. A change in the clearance angles leads to a change in the contact area of the rear faces of the tool with the machined surface. But, as experiments with tool wear showed, increasing the wear land to \(0.4\) mm is not reflected in the quality of the surface. Consequently, surface finish should not depend on the principal clearance angle of the tool. Indeed, as the authors’ experiments showed, changing the principal clearance angle \(\alpha\) from 8 to \(15^\circ\) produced no substantial change in the height of the micro-irregularities.

Fig. 7

The profile of the cutter depends on the magnitude of the plan angles \(\varphi\) and \(\varphi_1\). But in finishing operations, as indicated above, the entire profile of the micro-irregularities on both sides is formed by an arc of radius \(r\), and the geometry of the profile of the micro-irregularity does not depend on changes in the plan angles.

The principal geometrical characteristic of the profile of the micro-irregularity is the nose radius of the cutter \(r\). As can be seen from Table 1, increasing the nose radius \(r\) should lead to a considerable reduction in the heights of the micro-irregularities. At the same time, increasing the radius, owing to insufficient rigidity, leads to the appearance of vibrations and chatter.

which adversely affects surface finish. Indeed, S. P. Semenov [9] found in fine boring that increasing the nose radius initially leads to a decrease in the height of microroughness, and then the surface finish deteriorates. According to Semenov, the optimum nose radius, at which the height of microroughness is the smallest, is \(r = 1.5 \div 2.0\) mm. The same dependence \(H_{\max} = f(r)\), as can be seen from Fig. 8, was also obtained by the authors.

Fig. 8

In conclusion, it may be stated that in finishing operations the geometry of the tool (with the exception of the nose radius) should be selected on the basis of its optimum durability. The magnitude of the nose radius depends on the rigidity of the machine tool. The greater the rigidity of the machine tool, the larger the nose radius of the cutter should be.

Influence of Tool Wear on Surface Finish

The question of the influence of tool wear on surface finish is treated in detail in the works of P. E. Dyachenko and M. O. Yakobson [4] and [6]. In these works the authors come to the conclusion that the dependence of \(H_{\max}\) on wear along the flank face, \(h\), when cutting with high-speed-steel tools and with tools fitted with carbide inserts, is different. When cutting with high-speed-steel tools, the curve \(H_{\max} = f(h)\) has the same character for different materials: as wear increases, \(H_{\max}\) initially decreases and then increases. The authors of [4] and [6] explain this regularity by the fact that, as the tool becomes dull, the cutting pressure increases, since with an increase in the wear land on the tool flank the area of contact between the tool nose and the workpiece, and the friction of the tool against the workpiece, increase. In addition, they attempt to relate the curve \(H_{\max} = f(h)\) to the curve of chip-shrinkage dependence on wear. According to the same authors, when cutting with tools fitted with carbide inserts, the curve \(H_{\max} = f(h)\) changes very irregularly, which is likewise explained by the fact that, when the cutting edge becomes dull, small particles break out of it.

In all their experiments to determine the influence of wear of a carbide tool on surface finish, the authors of the present article found that the curve \(H_{\max} = f(h)\) has the same character as that obtained by Dyachenko and Yakobson in the case of cutting with high-speed-steel tools. That is, with an increase in \(h\), the value of \(H_{\max}\) initially decreases and then increases (Fig. 9). The authors believe that such a regularity cannot be explained by a change in chip shrinkage. As the works of Yu. A. Rozenberg [7] and A. K. Baikalov [11] show, it may be assumed that chip shrinkage does not depend on wear along the tool flank face. The regularity of the curve \(H_{\max} = f(h)\) should be explained by a change in the tool profile and by the nature of the change in temperature and cutting forces on the flank face during the process of tool wear. As a result of wear, the nose radius of the cutter increases and the auxiliary clearance angle decreases, which leads to a certain reduction in the height of microroughness. In addition,

In addition, it is interesting to note that the curves of the variation of the specific friction force and the specific normal force on the flank face with increasing wear land, obtained by A. K. Baikalov [11], have a character similar to the curve \(H_{\max}=f(h)\). A. K. Baikalov indicates that the cause of the initial wear of the tool is associated with large values of the specific friction force and the normal force on the flank face of the hard alloy. At a certain value of the wear land \((h>0.2\ \text{mm})\) the specific forces decrease and change little up to a wear of \(h=0.8\ \text{mm}\). Beginning from a certain point, these

Fig. 9

forces increase, which leads to an increase in the intensity of wear toward the end of dulling. In exactly the same way, one may assume a connection between the change in these specific friction and normal forces on the flank face of the tool and the surface quality.

In conclusion it should be noted that, as the authors’ experiments show, tool wear on the flank face up to a wear land of \(0.8\ \text{mm}\) does not worsen the finish class.

Conclusions

1. In turning gray cast iron on machine tools of medium rigidity, the economically attainable finish class is: for feeds \(S<0.07\ \text{mm/rev}\)—▽▽6a, and for feeds \(S=0.07 \div 0.3\ \text{mm/rev}\)—▽▽5.

2. In finishing operations the cutting conditions must correspond to a tool life \(T\leq 90\) minutes, because in this case the possibility of built-up edge formation on the tool is eliminated. The choice of feed in this case is determined by the required finish class and the rigidity of the machine tool.

3. In finishing operations the tool geometry, with the exception of the nose radius of the tool, should be selected on the basis of its optimum life, since the parameters of tool geometry under finishing conditions do not affect surface finish. The value of the optimum nose radius of the tool depends on the rigidity of the machine tool. The greater the rigidity of the machine tool, the larger the nose radius should be. When working on machine tools of medium rigidity, the optimum value of the tool nose radius is \(1.5\text{—}2.0\ \text{mm}\).

4. Tool wear on the flank face up to a wear land of \(0.8\ \text{mm}\) does not worsen the surface finish.

5. In cutting gray cast iron, as in cutting steels, the regularity of the change in unevenness as a function of cutting temperature is represented on the graph by a typical curve, the character of which does not depend on the factors that change the temperature.

References

1. Semenov S. P. Formulation of the problem of surface finish of machine parts and the history of its development in the USSR, LONITOMASH, book 16. Surface Quality of Machine Parts. Part II. Mashgiz, 1950.

2. Dolmatovskii G. A. Technologist’s Handbook. Mashgiz, 1950.

3. Beletskii D. G. Microgeometry of the Surface in Turning. Mashgiz, 1947.

4. Dyachenko P. E. and Yakobson M. O. Surface Quality in Machining Metals by Cutting. Mashgiz, 1951.

5. Eremin A. N. The Physical Nature of Phenomena in Cutting Steel. Mashgiz, 1951.

6. Dmychenko P. E. Investigation of the Dependence of Surface Microgeometry on Machining Conditions. Publishing House of the Academy of Sciences of the USSR, 1949.

7. Rozenberg Yu. A. Investigation of the Cutting Process of Gray Cast Iron. Dissertation, Tomsk, 1952.

8. Isaev A. I. The Process of Formation of the Surface Layer in Metal Cutting. Mashgiz, 1950.

9. Semenov S. P. Fine Boring of Cylinders and Sleeves. Mashgiz, 1949.

10. Rog V. A. Surface Finish of Parts in Automotive Repair Production. Publishing House of the Ministry of Communal Economy of the RSFSR. Moscow, 1954.

11. Baikalov A. K. Optimal Geometry of Cutters in High-Speed Turning of Gray Cast Iron. Dissertation, Tomsk, 1953.