PROCEEDINGS
OF THE TOMSK ORDER OF THE RED BANNER OF LABOR POLYTECHNIC
INSTITUTE named after S. M. KIROV

Volume 129                   1965

N. P. RYASHENTSEV, A. P. TRONOV, A. V. FROLOV, V. I. KOPYTOV

SELECTION OF THE STRIKER MATERIAL FOR ELECTROMAGNETIC HAMMERS

(Presented by the Department of Mining Machines and Mine Transport)

An analysis of the energy diagram of solenoid hammers with free striker travel [2] shows that the principal losses are losses in the copper of the coil and losses in the steel of the magnetic circuit, which together amount to as much as 60% of the total power consumed.

For the magnetic circuit of a hammer made of sheets of weakly alloyed steel E12, insulated from one another, the specific losses due to hysteresis and eddy currents at an induction of 14000 gauss are about 6 W/kg, that is, about 2% of the total power consumed. Consequently, the principal losses in solenoid hammers are ohmic losses in the copper of the coil and losses due to eddy currents in the striker. The ohmic losses of the hammer can be limited by reducing the active component of the coil current, the magnitude of which is proportional to the losses from eddy currents. By reducing eddy-current losses in the striker, the efficiency can be increased and the heating of the hammer reduced. The principal means of reducing eddy-current losses in the striker are:

1. The development of hammer power-supply circuits that make it possible to obtain a flat peak of the induction magnetization curve.

2. The use of ferromagnetic materials with low specific electrical conductivity and high magnetic properties.

3. The choice of the form and design of the striker, and of the method of its mechanical and heat treatment, ensuring minimal electrical losses per unit of useful work.

In solenoid hammers with free striker travel, it does not yet appear possible to change the power-supply circuit in order to reduce eddy-current losses. It is therefore necessary to seek other ways of reducing losses.

In existing designs of solenoid hammers of the types MS-15 m, MS-20, MS-18/36 m, the pole pieces and yoke are stamped from sheet electrical steel [3, 4], while the striker is made as a solid metal piece, since it must possess sufficient mechanical strength under a cyclic-impact mode of operation. During operation of the hammer, the striker strikes with its end face against the chuck into which the tool is inserted. In this connection, when selecting the material of the striker it is necessary to proceed from its resistance to impact-fatigue failure.

The wear resistance of the striker is not a criterion determining the choice of its material. This was confirmed by the results of laboratory and industrial tests of hammers with a guide tube made of stainless steel 1Kh18N9T (the inner surface of the tube was chromium-plated). When the hammer operated without lubrication for 100–150 hours, the wear was 0.01–0.02 mm, and the parts were suitable for further operation.

In addition, the striker must have minimal eddy-current and hysteresis losses when the hammer is supplied with pulsating current (according to a single-half-period rectification circuit). At the same time, the striker must be made of a material with high permeability in medium and strong magnetic fields and, in particular, with high saturation. The values of the initial and maximum permeability, as well as the residual induction and coercive force of the core, are of little significance for the striker material, although it is desirable that the last two quantities be minimal.

The electromagnetic properties of materials given in works [5–7] cannot be fully used for selecting the striker material for the following reasons: the magnetic flux passing through the striker is nonsinusoidal; magnetization reversal of the striker does not occur; the striker moves in a varying magnetic field and partially intersects leakage and stray fluxes.

We have carried out additional investigations to develop recommendations for selecting a striker material that satisfies the strength and electromagnetic requirements to the greatest extent.

Taking into account that the hammer is a fast-acting machine (cycle duration 40 msec), the striker should be made of a material with minimal eddy-current losses, which affect not only the heating of the hammer but also increase the time of striker engagement. In designing hammers, it is advisable to adopt a striker induction close to the saturation induction. This gives the minimum internal diameter of the coil, which leads to a reduction in ohmic losses.

The losses in the striker (W/kgf), obtained on the basis of the results of experimental investigations, are given in Table 1.

Table 1

It is seen from the table that the smallest losses are in a striker made of special cast iron. However, the striker material cannot be recommended on the basis of losses in steel alone, since the total losses also include losses in the copper of the coil, which depend on the striker material. In addition, cast iron does not satisfy the requirements of the cyclic-impact operating regime.

Of the steel grades considered (Table 1), it is seen that steel 9KhS most fully satisfies all operating conditions of the striker. A striker made from it, with appropriate heat treatment, provides sufficient mechanical strength and minimal electrical losses. Comparing the chemical composition and mechanical properties of steels 9KhS and 40KhS, we note that the carbon content in the latter grade

became 2 times smaller, while silicon became 25–30% greater. This gives grounds to assert that the losses in 40KhS steel will be smaller than in 9KhS. Thus the efficiency of a hammer operating with a striker made of 40KhS steel will be higher than that of a hammer with a striker made of 9KhS steel (because of the absence of 40KhS steel, comparative experiments were not carried out).

Above, a cylindrical striker without grooves and heat treatment was considered; these have great significance not only for the losses in the steel of the striker, but also for the total losses and the working process of the hammer, since the hammer coil is supplied with pulsating current obtained by a single-phase rectifier circuit. Consequently, the striker and the magnetic circuit are in a pulsating magnetic field. Under these conditions, eddy currents arise in the striker, creating their own magnetic flux, which is directed opposite to the main flux. The magnetic flux created by the eddy currents leads to a reduction of the resultant flux inside the striker (skin effect). Under real conditions the phenomenon of the skin effect leads to a change in the dynamic inductance.

The theoretical determination of the depth of penetration of the magnetic flux for the given operating conditions of the striker is difficult; therefore we solved this problem experimentally. For this purpose, experiments were carried out with a cylindrical striker made of cast iron

Fig. 1. Dependences of the specific consumed power \(p_a\) (W/kgf) and copper losses \(p_m\) (W/kgf) on induction and striker cross section. \(p_a^I, p_a^{II}, p_a^{III}, p_a^{IV}\) respectively indicate the active losses per unit induction for a hammer with a striker without a longitudinal hole and with one groove, and then with the same striker inside which through holes of diameter 6, 9.5, and 14 mm were drilled. \(p_m^I, p_m^{II}, p_m^{III}, p_m^{IV}\) are the copper losses per unit induction for a hammer with a striker without a hole and with one groove \((p_m^I)\), and with a striker inside which holes of 6, 9.5, and 14 mm, respectively, were drilled.

SCh-28-48, with one groove 8 mm deep. The dependences of the consumed power \(p_a\) per unit induction and the specific ohmic losses \(p_m\) were recorded as functions of the mean value of the maximum induction \(B_m\) of the transverse section of the striker. The results of the experimental data are shown in Fig. 1.

If the cross-sectional area of the striker is taken as 100%, then after drilling a longitudinal hole with diameters of 6 mm, then 9.5 and 14 mm, the cross section of the striker will respectively become: 95, 88, and 75%.

From the dependences given (Fig. 1) it is evident that, with increasing induction, the specific losses \(p_a\) and \(p_m\) increase, while with increasing diameter of the hole in the striker they decrease. This means that the ratio of induced eddy currents to induction in the striker decreases as the drilled hole is enlarged, i.e., the degree of utilization of the striker improves. Our task included not only determining the degree of utilization of the striker as an active material, but also revealing possibilities for increasing the efficiency of the hammer as a whole.

The specific active losses \(p_{a.\phi}\) and the specific losses in copper \(p_{m.\phi}\) per unit of flux were determined from the expression

\[ p_{a.\phi}=\frac{P_{a.\phi}}{\phi}; \qquad p_{m.\phi}=\frac{P_{m.\phi}}{\phi}, \]

where \(P_{a.\phi}\) is the active consumed power, W; \(P_{m.\phi}\) is the copper loss, W; \(\phi\) is the magnetic flux, klines.

Fig. 2. Dependences of the specific consumed power \(p_a\) (W/kline) and copper losses \(p_m\) (W/kline) on magnetic flux and striker cross section. \(p_{a.\phi}^{I}\), \(p_{a.\phi}^{II}\), \(p_{a.\phi}^{III}\), \(p_{a.\phi}^{IV}\) respectively indicate the specific active losses for a striker without a hole and for a striker with holes of 6, 9.5, and 14 mm. \(p_{m.\phi}^{I}\), \(p_{m.\phi}^{II}\), \(p_{m.\phi}^{III}\), \(p_{m.\phi}^{IV}\) are the specific copper losses under the same conditions.

From the obtained dependences (Fig. 2) it is evident that the losses \(p_{a.\phi}\) and \(p_{m.\phi}\) increase faster than the magnetic flux increases. The smallest losses \(p_{a.\phi}\) and \(p_{m.\phi}\) occur in the hammer with a striker without a hole. With increa-

with an increase in the drilled hole the losses increase. If the cross section of the striker is reduced by 12% when drilling the hole, then at \(\varphi = 90\ \text{km/sec}\) the losses \(p_{a.\phi}\) increase by 12.5%, and \(p_{m.\phi}\) by 17.5%. With an increase in the hole the losses increase still more.

A reduction of eddy currents in the striker can also be obtained by another method—by making longitudinal slots. To determine the influence of slots on the specific active losses \(p_a\), the specific losses in copper \(p_m\), and the specific losses in the striker steel \(p_{c.b}\), experiments were carried out and the dependences of the indicated losses on the magnitude of the induction (Fig. 3, a) and on the number of slots (Fig. 3, b–d) were plotted. With increasing induction, the specific losses \(p_a\), \(p_m\), and \(p_{c.b}\) increase. A single slot made in the striker

Fig. 3. Dependences of the specific consumed power \(p_a\) (W/kgf), losses in copper \(p_m\) (W/kgf), and losses in the striker \(p_{c.b}\) (W/kgf) on magnetic induction and the number of slots. \(p_a^1\), \(p_a^{II}\), \(p_a^{IV}\), \(p_a^{VIII}\), respectively, show the active losses per unit induction for a hammer with a striker on which 1, 2, 4, or 8 slots are made; \(p_{c.b}^1\), \(p_{c.b}^{II}\), \(p_{c.b}^{IV}\), \(p_{c.b}^{VI}\), respectively, show the losses in the striker steel per unit induction when 1, 2, 4, or 8 slots are made on the striker; \(p_m^1\), \(p_m^{II}\), \(p_m^{IV}\), \(p_m^{VIII}\), respectively, show the losses in the copper of a hammer with a striker on which 1, 2, 4, or 8 slots are made.

one slot significantly reduces the induction of eddy currents and limits the influence of the surface effect. This is evident from the fact that, when one slot is made in the striker, the losses \(p_a\) (at striker induction \(B_m = 14000\text{–}17000\ \text{Gs}\)) decrease by 15–25%; when two slots are made, the indicated losses are further reduced by 12–20%. With a further increase in the number of slots, the losses decrease. However, as the number of slots increases, the dependence \(p_a = f(B_m)\) becomes flatter (Fig. 3, d), and when the number of slots is increased from six to eight the reduction in losses is insignificant. The specific losses in copper \(p_m\) and in the striker steel

$p_{c.b}$ decrease as the number of slots increases. If the slot width of a 28 mm striker is increased from 1 to 2 mm, then the losses in the striker steel decrease in accordance with the reduction of its weight, the losses in the copper increase by 53–58%, and the active losses increase by 10–24%. Therefore, the slot width on the striker should be made minimal.

The decrease in losses $p_a$ when slots are present on the striker is explained by a reduction in the active component of the current passing through the coil. As a result, the ohmic losses $p_m$ in the coil, determined by both the reactive and active components of the current, also decrease. In this case the decrease in the active component of the current occurs more intensively than the increase in the reactive component. When more than eight slots are made, the losses $p_a$ and $p_m$ may increase as a result of a decrease in the cross-sectional area of the striker and an increase in the parasitic gaps between the pole and the striker. In this case the increase in the reactive component occurs more intensively than the decrease in the active component.

Fig. 4. Dependences of the consumed power $p_a$ (W/kgf), losses in copper $p_m$ (W/kgf), and losses in the striker $p_{c.b}$ (W/kgf) for a hammer with a striker made of electrical steel E-4 (a) and for steel 9KhS (b, c, d) on induction and heat treatment. $p_a^{\mathrm{VIII}}$ is the specific consumed power of a hammer with a striker of steel 9KhS without heat treatment; $p_a^{0}$—with an annealed striker; $p_a^{T}$—with the striker faces hardened to hardness $HRC = 47 \div 49$; $p_a^{n}$—with a striker hardened to hardness $HRC = 57 \div 59$ followed by tempering. The corresponding indices denote the losses in the coil copper $p_m$ and in the striker steel $p_{c.b}$.

In practice, in order to reduce electrical losses, magnetic cores operating in alternating and pulsating magnetic fields are made laminated, from sheet steel with improved magnetic characteristics. We manufactured a striker wound from sheets of E-4 steel (Fig. 4, a), previously coated with BF-2 glue.

Then the specific consumed power \(p_a\), the specific losses in copper \(p_m\), and in steel \(p_{c.b}\) were determined for the hammer with this striker. Comparing the indicated losses of a hammer with a twisted and solid-metal striker with grooves (\(p_a^{\mathrm{VIII}}\), \(p_m^{\mathrm{VIII}}\), and \(p_{c.b}^{\mathrm{VIII}}\)), determined for a hammer with a thermally untreated striker made of 9KhS (Fig. 4, 6), it was established that the losses in the steel of the twisted striker \(p_{c.b}\) are almost 2 times smaller than in the 9KhS striker, while the losses in copper \(p_m\) increased by more than 2 times. Overall, the specific consumed power of the hammer with the twisted striker increased by a factor of 1.1–1.7 in comparison with the consumed power of the hammer with a striker made of 9KhS steel. In addition, the technology for manufacturing a twisted striker is much more complicated than the technology for manufacturing a striker from 9KhS steel. Therefore, it is inexpedient to manufacture a twisted striker of the design described.

To determine the influence of heat treatment on losses, experiments were carried out with a striker without heat treatment, annealed, with hardened end faces, and with a striker hardened over the entire surface with subsequent tempering.

From the dependences presented (Fig. 4), it is seen that the minimum losses \(p_a^\circ\), \(p_m^\circ\) occur in the striker with full annealing. This was to be expected, since annealing removes work hardening and promotes the formation of coarse grains, and an increase in grain size is considered one of the effective means of improving the magnetic properties of steel and reducing eddy-current and hysteresis losses [5–7]. With increasing induction, the specific losses \(p_a^\circ\) and \(p_m^\circ\) increase faster than the induction, while \(p_{c.b}^\circ\) increases almost in proportion to the increase in induction. A large increase in \(p_{c.b}^{\mathrm{VIII}}\) is observed at \(B_m = 17000\text{–}18000\) G.

The losses \(p_a^{T}\), \(p_m^{T}\), \(p_{c.b}^{T}\) for the striker with hardened end faces are greater than the losses \(p_a^{\mathrm{VIII}}\), \(p_m^{\mathrm{VIII}}\), \(p_{c.b}^{\mathrm{VIII}}\) for the striker without heat treatment, and still greater than \(p_a^\circ\), \(p_m^\circ\), \(p_{c.b}^{\mathrm{VIII}}\) for the striker with full annealing. When the striker is hardened over the entire surface to a hardness of \(H_{Rc} = 57 \div 59\), the losses \(p_a^n\), \(p_m^n\), and \(p_{c.b}^n\) are maximal. It should be noted that the specific losses of a hardened striker without subsequent tempering are much greater than those of a striker hardened to the same hardness but with full tempering. This is understandable, since during tempering the magnetic properties of the steel that it had before hardening are restored to some extent.

Conclusion

1. When selecting the material for the striker, it is necessary to take into account the cyclic-impact mode of its operation and to choose a material with a minimum amount of harmful impurities. After heat treatment, internal stresses should be removed while at the same time obtaining optimal grain sizes that ensure improvement of the magnetic properties of the steel and minimal eddy-current and hysteresis losses.

2. The strikers of electromagnetic hammers should be made from 9KhS steel, which provides minimal specific losses in copper and steel at the minimum consumed power per unit of striker induction in medium and strong magnetic fields.

3. Drilling a longitudinal hole in the striker reduces the influence of the surface effect and makes it possible to use the striker better as a magnetic core. However, the specific losses of a hammer with a striker in the center of which a hole has been drilled are greater than those of a hammer with a striker without a hole. Therefore, it is inexpedient to make holes in the strikers of light hammers.

1. Longitudinal grooves made in the striker reduce the induction of eddy currents and lessen the influence of the surface effect. For a striker 28 mm in diameter, made of 9KhS steel, it is advisable to make 6–8 grooves 1 mm wide and 7–8 mm deep. In this case the specific losses in the copper of the coil and in the steel of the striker, as well as the power consumed per unit of induction (at \(B_m = 14\text{–}15\ \mathrm{kGs}\)), are reduced by a factor of 2–2.5. A larger number of grooves should not be made for a striker of this diameter, since this will lead to a reduction in the hammer efficiency.

2. A striker made of 9KhS steel must be annealed; then only the end faces of the striker should be hardened, followed by tempering. After such heat treatment the striker has sufficient mechanical strength under a cyclic-impact mode of operation and ensures minimal specific losses.

REFERENCES

1. N. P. Ryashentsev, A. V. Ivanov. Solenoid hammers for installation work. Technical-Economic Bulletin of the Kemerovo Sovnarkhoz, No. 10, 1960.

2. N. P. Ryashentsev, E. M. Timoshenko. Experimental investigation of the working process and energetics of a solenoid hammer with free striker travel. Izvestiya TPI, Vol. 123, TGU Press, 1963.

3. N. P. Ryashentsev, O. D. Alimov, E. M. Timoshenko, A. V. Frolov. Solenoid hammer. Inventor’s certificate No. 132578 of February 18, 1960.

4. N. P. Ryashentsev, A. V. Frolov, E. M. Timoshenko. Magnetic core of a solenoid hammer. Inventor’s certificate No. 139622 of December 16, 1960.

5. A. S. Zaimovskii, L. A. Chudnovskaya. Magnetic Materials. Gosenergoizdat, 1957.

6. A. S. Zaimovskii, V. V. Usov. Metals and Alloys in Electrical Engineering. Gosenergoizdat, 1941.

7. B. G. Livshits. Physical Properties of Metals and Alloys. Mashgiz, 1956.