SIMULATION OF PNEUMATIC HAMMER VIBRATION DAMPING BY CHANGING THE SHAPES OF PRESSURE DIAGRAMS

A. V. Trikhanov, V. I. Baburov, V. F. Gorbunov

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

One of the principal sources of vibration in pneumatic hammers is the action of variable air pressure on the walls of the housing. To reduce the amplitude of oscillations of the housing, B. V. Sudnishnikov proposed changing the shape of the pressure diagrams [1, 2]. This proposal was implemented by a number of researchers.

Tests of the machines created showed that a reduction in the vibration amplitude of the hammer housing is achieved by a factor of 1.2–5 [3–5]. In our works [6–8] it is shown that, depending on the magnitude of the pressing force on the hammer housing, in addition to the internal disturbing force, other forces also act—for example, the forces of collision with the collar of the tool, the forces of the striker impact on the housing, etc. In this connection, the effect of vibration damping of the machine by changing the shapes of the pressure diagrams is also determined by the operating mode of the hammer.

The present work gives the results of identifying the quantitative possibilities for reducing vibrations of KE-type hammers by changing the shapes of the pressure diagrams above and below the striker. The investigations were carried out for an air-line pressure of 5 atm and for various pressing forces during work on steel. The machine under investigation (the KE-22 hammer) was positioned horizontally. To solve the problem posed, we selected five varieties of diagrams of working cycles (Fig. 1). The first cycle is the ordinary cycle of the KE-22 hammer. The fourth cycle is called optimal, since, according to data [2, 7], a machine operating according to this cycle will have the smallest vibration amplitude. The intermediate position between the indicated cycles is occupied by the second, third, and fifth cycles (Fig. 1), implemented by various researchers [3–5].

The study was carried out on an electronic modeling installation of the IPT-5 type, which made it possible to avoid labor-intensive work on the design, manufacture, and testing of experimental hammer prototypes. In addition, modeling makes it possible to obtain the vibration characteristics of hammers both with allowance for rebound impacts of the tool on the hammer housing and without them. In the latter case, the effect of vibration damping due to changing the shapes of the pressure diagrams can be revealed, so to speak, in pure form, which is impossible to obtain on experimental hammer prototypes.

In studying the five indicated cycles, the basic parameters of the hammer corresponding to the KE-22 hammer remained unchanged. Figure 1 gives the indicator diagrams of the cycles and the paths of the striker for the case in which reverse impacts of the tool are absent and the body is in constant contact with the tool collar.

Fig. 1. Diagrams of air pressures in the cavities of the hammer and displacement of the striker:
1 — normal cycle; 2 — second cycle; 3 — third cycle;
4 — “optimal” cycle; 5 — fifth cycle

The technical data of the hammer and the characteristics of the cycles are given in Tables 1 and 2.

It is seen from the tables that the calculated values and those obtained on the electronic model differ by no more than 11%. These values are close to the technical performance of the KE-22 hammer. Significant discrepancies occur only for the pressing force. The values of the pressing force of the actual KE-22 hammer are 4 kg higher than the corresponding values obtained on the electronic model [5].

By modeling, dependences were obtained for the vibration amplitude of the hammer body on the pressing force for various forms of pressure diagrams, both in the presence of reverse impacts of the tool against the body (Fig. 2) and without them (Fig. 3).

Let us consider the case in which reverse impacts of the tool against the body occur (Fig. 2). The character of the change in the vibration amplitude of the body as a function of the pressing force is the same for all forms. At the minimum pressing forces the body has the greatest vibration amplitude. In the floating mode the vibration amplitude of the body decreases to a certain value and again increases in the limiting basic mode.

Table 1

Then, as the pressing force increases, a sharp decrease in amplitude occurs, after which it again, despite the increase in pressing force, grows and then gradually decreases. Consequently, the dependence between vibration amplitude and pressing force (the amplitude–force dependence) has two maxima. This phenomenon was first discovered by us and explained in works [6, 7].

During electronic modeling, oscillograms were obtained for the motion of the body \(x_k\) and of the tool \(x_i\) for all five cycles. These oscillograms confirm the authors’ explanations of the second maximum of the amplitude–force characteristic.

Table 2

The oscillograms shown in Fig. 4 for one of the cycles illustrate this maximum well.

At pressing forces greater than 30 kg, the amplitude–force relationship for all forms differs within the limits of the errors of electronic modeling (both qualitatively and quantitatively). This is explained by the fact that the internal disturbing force acting

Fig. 2. Amplitude–force relationships of the hammer in the presence of rebound impacts: 1 — ordinary cycle; 2 — second cycle; 3 — third cycle; 4 — “optimal” cycle; 5 — fifth cycle

on the housing is equal to or less than the pressing force applied to the hammer handle.

The influence of rebound impacts in this case is the same, since the striker velocity before impact is kept constant. The impact interaction of the tool, the rock being worked, and the housing, naturally, also does not change.

Fig. 3. Amplitude–force relationships of the hammer in the absence of rebound impacts: 1 — ordinary cycle; 2 — second cycle; 3 — third cycle; 4 — “optimal” cycle; 5 — fifth cycle

A peculiarity is present only for the “optimal” cycle. For it, the second minimum turned out to be shifted to the left, i.e., to a smaller pressing force. The leftward shift of the second minimum for this relationship should be explained by the weak influence of the internal disturbing force. The influence of rebound impacts already begins to predominate after 22 kg, whereas for the other forms it does so after 27–29 kg.

From what has been considered, it may be concluded that at pressing forces greater than 25 kg the influence of the form is not substantial. In the marginal principal mode the ordinary cycle gives a vibration amplitude of 3.3 mm, the fifth 3.0 mm, the second 2.8 mm, the third 2.6 mm, and the optimal cycle 1.9 mm.

This means that, in the marginal principal mode, in comparison with the ordinary cycle, the fifth cycle damps the vibration amplitude by a factor of 1.1, the second by a factor of 1.2, the third by a factor of 1.3, and the “optimal” by a factor of 1.7.

Depending on the shape of the indicator diagram, the pressing force \(R_{\mathrm{n}}\) changes, ensuring the limiting basic mode. Thus, for the ordinary cycle \(R_{\mathrm{n}} = 21\) kg, for the second—19 kg, for the third—18 kg, and for the “optimal” cycle—17 kg. This is explained by the fact that, as the vibration amplitude decreases, the velocity of the housing at the moment of its seating on the tool collar decreases [1, 3, 5, 7]. This velocity enters into the expression for determining the pressing force of the limiting basic mode, which is given in [5, 7].

In the “floating” mode the vibration amplitude for the ordinary cycle is 2.72 mm, for the fifth—2.65 mm, for the second—2.48 mm, for the third—2.16 mm, and for the “optimal” cycle—1.76 mm.

This means that, in comparison with the ordinary cycle in the “floating” mode, the fifth cycle produces a vibration amplitude 1.02 times smaller; the second cycle—1.1 times smaller; the third cycle—1.26 times smaller; and the “optimal cycle”—1.54 times smaller. The pressing force ensuring the “floating” mode is 16 kg for all forms.

Fig. 4. Diagrams of oscillations of the housing of a pneumatic hammer and of the tool at pressing forces on the handle:
a) 50 kg; b) 35 kg; c) 27.5 kg

This confirms the conclusions drawn in [5], where it is stated that the shape of the pressure diagrams does not affect the magnitude of the pressing force that ensures the “floating” mode of operation of the hammer.

A comparison of the effectiveness of damping body vibration in “floating” and near-main regimes indicates that, in the near-main regime, this effect is somewhat greater. In general, the smaller the pressing force, beginning with the main regime, the smaller this effect.

In fact, at \(R_{\mathrm{p}} = 12\) kg almost all the dependencies already practically coincide.

This should apparently also occur in the case when there are no rebound impacts of the tool against the body. The amplitude–force characteristics in this case, too, independently of the forms of the indicator diagrams, have the same character of variation (Fig. 3). They are similar to the characteristics for the case when rebound impacts of the tool against the body are present (Fig. 2). The difference consists mainly in the fact that, at pressing forces exceeding 30–33 kg, body vibration is absent.

At pressing forces exceeding the pressing forces that ensure the near-main regime, the form of the indicator diagram substantially influences the vibration amplitude. Let us take a pressing force equal to 20 kg. With the usual cycle the vibration amplitude is 3.07 mm; with the fifth, 2.84 mm; with the second, 2.76 mm; with the third, 2.20 mm; and with the “optimal” cycle, 1.25 mm.

This means that, compared with the usual cycle, the fifth cycle provides vibration damping by a factor of 1.08, the second by a factor of 1.11, the third by a factor of 1.4, and the “optimal” cycle by a factor of 2.5.

At a pressing force of 25 kg, compared with the usual cycle, the fifth cycle provides vibration damping by a factor of 1.2, the third by a factor of 2.45, and the “optimal” cycle by a factor of 3.8.

Such an influence of the form of the indicator diagrams on the vibration amplitude of the body at the considered pressing forces is not observed for the case when the tool is movable. This is explained by the fact that, in these regimes, the influence of rebound impacts of the tool on the body already begins to manifest itself, which also reduces the effect of vibration damping due to a change in form.

Conclusions

1. A change in the form of the indicator diagrams of a pneumatic hammer, while its main parameters are preserved, affects the amplitude–force characteristic of the hammer. The influence of the form is most substantial for near-main operating regimes in the absence of rebound impacts. The vibration amplitude is reduced by as much as a factor of 3.8. At pressing forces greater than 30–35 kg, vibration is absent if there are no rebound impacts of the tool against the body.

2. When rebound impacts of the tool against the body are present, changing the cycle leads to almost no change in the amplitude–force characteristic for the main operating regimes of the hammer.

3. In any case, in the near-main regime it is possible to reduce the vibration of the hammer body by a factor of 1.7, and in the “floating” regime by a factor of 1.5.

4. The smaller the pressing force, beginning with the near-main regime, the smaller the damping effect.

5. The pressing force that ensures the “floating” regime does not change when the cycle is changed.

6. The pressing force that ensures the near-main regime decreases for the following sequence of cycles: usual, fifth cycle, second cycle, third cycle, “optimal” cycle.

7. The greatest vibration is produced by the usual cycle, and the least by the “optimal” cycle; therefore, in a pneumatic hammer one should implement a cycle close to the “optimal” one and reduce the influence of rebound impacts of the tool against the body.

References

1. B. V. Sudnishnikov. Some Problems in the Theory of Impact-Action Machines, Publishing House of the West Siberian Branch of the USSR Academy of Sciences, Novosibirsk, 1949.

2. B. V. Sudnishnikov. The influence of forward-stroke parameters on the power and recoil of a pneumatic hammer. In: Percussion-Rotary Drilling, Novosibirsk, 1956.

3. N. A. Klyushin. Pneumatic hammers with a new cycle reducing recoil. Candidate dissertation. Novosibirsk, 1958.

4. A. A. Goppen. Reduction of vibrations of light riveting hammers. Mashinostroitel, no. 8, 1963.

5. A. M. Petreev. On the theory of the working cycle of impact-action pneumatic machines. In: Mathematical Methods in Mining (conference proceedings), part I. Novosibirsk, 1963.

6. V. F. Gorbunov, V. I. Baburov, L. S. Redutinsky, A. V. Trikhanov, G. G. Pestov. Study of the vibration of a hand-held pneumatic hammer when working on various materials. Article in the present collection.

7. V. F. Gorbunov, V. I. Baburov. On the dynamics of the forces acting on the body of a pneumatic hammer. Article in the present collection.

8. V. I. Baburov, V. F. Gorbunov, A. V. Trikhanov. Causes of vibration of hand-held pneumatic hammers and methods of combating it. Article in the present collection.

9. B. V. Sudnishnikov. A theorem on the displacement of a mass during the action of a force. Mechanization of Mining Operations. West Siberian Branch of the USSR Academy of Sciences, issue 8, 1950.