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Thomas D. Barkand and William J. Helfrich
U.S. Department of Labor
Mine Safety and Health Administration
Cochrans Mill Road, P.O. Box 18233
Pittsburgh, Pennsylvania 15236


Mine hoists and elevators have experienced accidents with the potential for injuring or killing numerous miners. These accidents occurred on counterweighted hoisting systems when the mechanical brake failed while the cage was empty. This allowed the counterweight to fall to the bottom of the shaft, causing the cage to overspeed and crash into the headframe. Direct-current hoist motors have the potential for preventing this type of accident by incorporating a passive electrical braking system known as dynamic braking. Installation of dynamic braking required minimal control-system modifications and modest expense. The various circuit designs that can be configured to implement dynamic braking on a dc mine hoisting system are discussed. Selection of the load resistance for the desired dynamic braking system performance will be addressed. In addition, analysis of data obtained from tests conducted on a mine hoisting system, which was modified in the field to incorporate dynamic braking, will be discussed to provide technical credibility to the dynamic braking concept.


Mine hoists and elevators possess several safety features to prevent the man cage from crashing into the bottom of the shaftway. Safety catches can prevent this type of accident from occurring when the brakes or wire ropes holding the cage fail. Safety catches, however, are not normally installed on the counterweight to prevent the cage from crashing into the headframe.

Many hoisting systems rely solely upon the mechanical brakes to stop the cage when an emergency condition occurs. Under normal operation, the mechanical brakes are called upon only to hold the cage at a stopped position while the electrical drive equipment is used to control the speed of the hoisting system. The frequency with which the mechanical brakes are exercised is minimal when compared to the constant usage of the electrical drive equipment. However, when an emergency condition is detected, the majority of hoists in the United States rely solely upon the mechanical brakes to stop the hoist. The assumption is that the electrical drive equipment is inoperative and the mechanical brakes are 100 percent reliable.

Accident history has proven that this is not always the case.[1][2] These accidents have occurred when the emergency stop button was pushed, which defeated the retarding effort of the hoist motor, when the mechanical brakes were inoperative. This allowed the overhauling load to free-fall, with the final speed limited only by inertia and frictional forces. The high speed crashes at the final travel limit caused extensive mechanical damage and fatal injuries.

Hoisting systems (i.e. elevators and hoists), driven by direct-current (dc) motors possess the capability to prevent the cage from wrecking into the headframe or the shaft bottom. The hoist electrical drive and control system can be designed to limit the speed of the falling overhauling load. This electrical source of braking can be used to retard the speed of the cage when the mechanical brakes fail. It is commonly referred to as dynamic braking, or dynamic retardation.

Dynamic braking utilizes the ability of the direct-current drive motor to act as a generator. It can be connected to an electrical load such as a resistor. Thus mechanical energy can be converted to electrical energy and then be dissipated as heat. This concept can be adapted to a mine hoisting system to provide an electrical backup to the mechanical braking system. The kinetic energy of the falling overhauling load can be utilized to drive the motor. The motor would then generate electricity, which could be dissipated as heat in a resistor connected to the motor. Torque is required for the motor to generate electricity. The retarding torque will limit the speed of the falling overhauling load. The amount of retardation and the final speed of the cage would be dependent upon the motor terminal characteristics and the ohmic value of the dynamic braking resistor (Rdb). By proper design, a failsafe mechanical and electrical braking system can be configured, which would drastically reduce the possibility of a hoist cage wrecking into the shaftway bottom or the headframe.

This paper will explore the feasibility, application, and testing of dynamic braking as it applies to mine hoisting systems. Also included is a case-study analysis of a dynamic braking system that was field installed and subsequently tested on a mine hoisting system.


The basic dc motor loop circuit consists of a dc motor armature in series with a dc power supply. This power supply can be either a generator or an SCR bridge. The power supply converts the incoming AC voltage into a variable dc voltage, which is used to control the speed of the dc motor. The field of the motor is normally supplied from separate source that is either fixed (i.e. constant potential) or variable (i.e. field weakening). The field of the generator is also supplied from a separate source that is varied to control the speed of the hoist motor.

Any shunt-wound dc motor can operate as either a motor or generator. It is operating as a motor when the torque it produces is in the same direction as shaft rotation. Under these conditions power supplied by the ac line is consumed in the form of mechanical energy at the motor shaft output. When motor torque is in the opposite direction as shaft rotation, the motor is operating as a generator. This occurs when a mechanical load overhauls motor torque, thus reversing shaft rotation. Under these conditions power is supplied by the overhauling mechanical load, changed into electrical power by the dc generator (i.e. the hoisting motor), and delivered back to the ac power system. This is called regenerative braking, since power is actively generated into the ac power system instead of simply being dissipated by a resistive load.

The system is in the powering mode when motor action is taking place. It is in the inverting mode when generator action is taking place. If raising the cage produces positive shaft rotation, then the four quadrants of operation are defined as shown in figure 1. The motor torque and direction are directly proportional to the armature current and voltage, respectively.

Fig. 1. Four Quadrants of Operation

Unbalanced hoisting systems will operate in quadrants 1 and 4 for constant-speed conditions. Quadrant 1 represents the condition wherein the motor speed and motor torque are in the same direction. Thus the motor supplies a positive motive force to the cage. Quadrant 4 represents the condition of negative direction and positive torque, therefore, the motor is developing braking force. During deceleration and acceleration the hoisting system will operate in quadrants 2 and 3, respectively.

Counterweighted (balanced) hoisting systems will operate in all four quadrants of operation for constant-speed conditions. When the counterweight is heavier than the cage, the hoisting system operates primarily in quadrants 2 and 3. The motor acts as a generator for part of the hoist cycle and as a motor for the other part of the hoist cycle, depending on the cage load.

The ability of the motor to regenerate into the ac power system is a valuable characteristic which allows the motor to provide an electrical (regenerative) braking force. Under normal operation the motor control circuit functions to provide both motive power and braking power. The mechanical brakes are called upon only to provide a very low-speed stop of the hoist at the top or bottom of the shaftway or, in case of an emergency, to provide complete stopping of the hoist. The mechanical brakes are called upon very infrequently to completely stop the hoist; however, when they are called upon they must provide 100 percent of the stopping force. This places a severe burden on the mechanical brakes at the critical time, when it is determined that the hoist is expected to wreck if it is not stopped.

Normally the motor is of such construction and reliability that it would not experience a debilitating failure. However, failures do occur to the motor control circuits and the related power supply. These failures can cause the motor to lose all ability to provide both motive power and braking power. For instance, if the field supply is disabled or fails, the motor will not be able to develop any torque to continue motoring or regenerating. Likewise, if a failure occurs in the power supply for the motor, the motor will not be able to provide motive or regenerative power. These problems occur due to various failures in the motor control circuit. Additionally, if utility power is lost, the motor will not normally be able to provide motive or regenerative power.

Taking into account all of the above factors, in addition to the accidents that have occurred due to brake failures, a system that uses the retarding capabilities of the hoist motor is warranted. These are several variations of the dynamic-braking control scheme. However, the general philosophy is to allow the motor to dissipate the mechanical energy by generating electrical energy and dissipating the energy as heat in a resistor.

The fundamental idea is to size a resistor at a low ohmic value (about one ohm) and allow the resistor to be applied across the motor armature when armature power is lost to the motor. The motor field can either be separately excited or can incorporate a field loss circuit, which would connect the field across the motor armature and would therefore be self-excited. The latter method is favored since it would provide protection under situations of complete power loss. The basic control philosophy can be divided into three categories: dynamic braking separately excited, dynamic braking self-excited, and dynamic braking separately or self-excited. The following is a description of each of these methods.

Separately Excited Dynamic Braking

Separately excited dynamic braking is the simplest form of dynamic braking to incorporate into a system; however, it only provides dynamic braking when the motor armature supply fails. It is assumed that under this condition the field power supply is functioning correctly and will continue to provide current to the motor field. This method of dynamic braking can be incorporated by installing a separate normally closed contactor for the dynamic braking resistor. A double-acting contactor could also be utilized, which provides a normally open (NO) contact for the motor loop and a normally closed (NC) contact for the dynamic braking resistor. The normally closed contact connects the dynamic braking resistor in parallel with the motor armature as shown in figure 2. The system configured in this method will provide dynamic braking any time the motor armature power supply fails or a fault occurs which initiates the opening of the loop contactor (M). In the event of a motor field supply failure, this system would not provide dynamic braking.

Fig. 2. Separately Excited Dynamic Braking System

Self-Excited Dynamic Braking

Self-excited dynamic braking is more difficult to incorporate than the previous method since it requires additional contacts to allow the field to be energized from the dynamic braking resistor. Figure 3 is an illustration of this control configuration. Every time the loop contactor opens, the field would be automatically shunted across the dynamic braking resistor. The field current would be supplied from the hoisting motor, generating into the dynamic braking resistor in parallel with the field. As the hoisting motor generates higher voltages, the field excitation would increase, and as a result the motor armature voltage would increase. Thus the motor becomes self-excited and requires no external power source for the motor field. This method is preferred over the first method since it requires no external power supply to provide dynamic braking and would function to prevent the hoist from catastrophically wrecking when there is a total power loss at the mine and the mechanical brakes fail. This system, however, is a little more complicated than the previous system. The complexity is offset by its ability to function under the most adverse condition of total power system failure.

Fig. 3. Self-Excited Dynamic Braking System

Separately or Self-Excited Dynamic Braking

This method of dynamic braking is the most complicated to design and expensive to install when compared to the previous two methods. However, it does provide the most desirable braking of the hoisting system. This control scheme is illustrated in figure 4. The method requires a device to sense when there is no field current being supplied by the motor field power source. This is usually accomplished by using a field-loss relay that is held closed as a function of the current in the field. When field current is lost, the relay drops out.

Fig. 4. Separately or Self-Excited Dynamic Braking

By using the field supply when it is available, the hoist is held at a relatively constant speed over the complete hoisting cycle. Therefore, the hoist does not accelerate into an overspeed condition before dynamic braking effort is developed, as it does when the self-excited form of dynamic braking is employed. This method, however, does give the advantage of having dynamic braking available when no power is available at the mine. It combines the best control characteristics of the previous two methods.


This review of dc machine characteristics is presented to provide a basis for analyzing the performance of the previously discussed dynamic braking circuit configurations. The generated voltage and torque relationships for dc motors and generators are summarized in the following equations:

eg = Kif W     V (1)
T = Kif ia N � m (2)
    eg generated voltage (V),
    if field current (A),
    W angular velocity (rad/s),
    T torque (N.m),
    ia armature current (A),
    K proportionality constant.
These equations assume that the magnetic system of the machine is linear. This is true when magnetic saturation, hysteresis, and magnetic retention in the field poles are neglected. The magnetization curve illustrating the effect of saturation but neglecting hysteresis and magnetic retention is shown in figure 5.

Figure 5. Magnetization Curve

The equations are valid where the slope of the line, K, is constant. The graph shows that the proportionality constant, K, is a measure of both the torque per ampere of armature current and the generated voltage (eg/W ) per ampere of field current.

The constant K can be determined from the manufacturers torque/armature current curves for motors or generated voltage/current curves (at rated speed) for generators. If these curves are not readily available, the constant can be approximated from nameplate information and empirical data. Nameplate data for motors will typically include rated horsepower, rated armature current, rated armature terminal voltage, rated speed, rated field current, and rated field resistance. The nameplate data for generators is similar, except the power rating is given to kilowatts.

The generated voltage, eg, is related to the terminal voltage by (3). The minus signs define the voltage drop for motors and the plus signs define the voltage drop for generators:

eg = V � ia Ra � 2     V (3)


    V g terminal voltage,
    Ra armature resistance,
    2 brush voltage drop.
The armature resistance, Ra, is usually not included on the nameplate, but it is needed to accurately calculate the value of K. The armature resistance can be obtained from the manufacturer or from empirical data.

The armature resistance varies nonlinearly from 0.15 to 0.08 ohms for 50-200 hp motors, respectively [3]. This low-resistance value can be accurately determined with a Kelvin bridge.

Since this type of instrument is not usually available at the mine, the resistance can be determined experimentally by measuring the armature voltage and current during a locked-rotor stall-current test. The magnitudes of armature voltage and current can be measured during testing with the voltmeters and ammeters typically used at the mine. The armature resistance is calculated from the armature current and corresponding voltage drop using Ohm's law.

Once the armature resistance is determined, the proportionality constant K can be calculated by substituting (3) into 91) and solving for K:

K = (Vr � iar Ra � 2)   -   (ifr Wr) (4)


    Vr rated armature voltage (V),
    iar rated armature current (A),
    ifr rated field current (A),
    r rated angular velocity (rad/s).
Error is introduced in the calculations if the rated field current is in the saturated region of the magnetization curve. However, this error should be small since the field excitation currents are typically not rated at a high degree of saturation. Once the K constant is known, the motor characteristics can be used to analyze the performance of the hoisting system under dynamic braking conditions.


When dynamic braking is activated the hoist motor operates as a generator. Generators can be classified by the means used to provide excitation for the field windings. They are either separately excited or self-excited. The dynamic braking system logic may require both methods of excitation to be used for different situations. For example, the control configurations shown in figures 3 and 4 will undergo a transition from separately excited to self-excited power generation when dynamic braking is activated.

Power generation characteristics of generators are typically analyzed at rated speed. This greatly simplifies load design calculations. However, designing dynamic braking performance requires analysis of power generation characteristics over a wide range of operating speeds, from 125 to 25 percent of rated speed. This adds an additional variable in the system design calculations.

The mechanical load on hoisting systems can vary from no load to full load. The dynamic braking system should be designed to safely lower the maximum overhauling load. However, the minimum load should not be decelerated faster than 16 ft/s2 (0.5 g force) during compound braking conditions. Compound braking effort is produced by the simultaneous activation of the mechanical and dynamic brakes. Compound braking occurs if the dynamic braking logic is designed to always back up the mechanical brake. Therefore, during emergency stopping conditions, the braking deceleration is produced by the summation of the dynamic and mechanical braking forces if the mechanical brake is operative.

The power rating of the dynamic braking resistor should be sized to dissipate the power generated when lowering the maximum overhauling load through the full length of the mine shaft under steady-state conditions. In addition, the dynamic braking resistor must be capable of dissipating the energy stored in the inertia of the hoisting system and transient electrical power generated during deceleration and switching conditions.

The sizing of the dynamic braking resistor will be analyzed in the following, based on the two methods of field excitation with respect to the speed, load, and power requirements for steady-state conditions. The transient response will be addressed later in the case study.

Separately Excited Dynamic Braking

A separately excited dynamic braking system is one in which the source of the field current is external to the machine. Connections for a separately excited dynamic braking configuration are shown in figure 6.

Figure 6. Separately Excited Dynamic Braking System

For steady-state conditions where if, , and ia are constant, the armature terminal characteristics are described in the linear range of the magnetization curve by the following equations:

Va = Kif s - Ra ia     V (5)
ia = (Kif W)   -   (Ra + Rdb) A. (6)

Generators are typically driven at rated speed by an ac motor. However, the inherent dynamic braking design requires the final dynamic braking speed of the hoist motor/generator to be less than rated speed. In addition, the final dynamic braking speed is dependent on the variable cage load. Therefore the dynamic braking system must operate over a speed range necessary to generate retarding torque from minimum to maximum load conditions. The effect of speed on the magnetization curve is shown in figure 7. For a given field current, the generated voltage is shown to be directly proportional to the speed as described earlier by (1). The typical terminal characteristics of the armature circuit (5) are shown in figure 8.

Figure 7. Separately Excited Magnetization Curves
for Various Percentages of Rated Speed

Figure 8. Terminal Characteristics of Separately Excited Generator
(Constant Field Current) for 100, 50, and 25 Percent of Rated Speed

The terminal voltage of a separately excited generator drops linearly with an increase in load current because of the armature circuit resistance. At high values of armature current an additional drop in terminal voltage may occur as shown in figure 8. This results from a reduction in the air-gap flux caused by the nonuniform saturating effect of the cross-magnetizing armature reaction.[5]

The cross-magnetizing saturation effect can be minimized if the magneto motive force of the field winding exerts predominate control on the air-gap flux, so that the condition of weak-field and strong armature magneto motive force may be avoided. The tendency toward distortion of the air-gap flux distribution also occurs when the field excitation remains substantially constant while the armature current exceeds rated values during heavy loads or transient conditions. However, this effect can often be neglected for armature currents below the rated value.

Under normal operating conditions the hoisting machine also operates as a motor. Figure 8 shows the terminal characteristic for reverse (motoring) armature current ia. High values of armature current reduce the flux and generated voltage as a result of the saturating effect of the cross-magnetizing armature reaction mentioned earlier.

To select the proper ohmic value for the dynamic braking resistor, the required retarding torque at the desired braking speed must be determined. The required braking torque is load-dependent. The cage loads can vary substantially from no load to full load. The worst-case loading of the hoisting system is typically full load. This provides the maximum inertia and the greatest imbalance. The greatest imbalance occurs at full load because the counterweight typically weighs 40-45 percent of the rated cage capacity (i.e. full load).[3] A 40-45 percent counterweighting provides the most power efficient operations since the hoisting system are typically utilized for light loads such as man-hoisting. Therefore, if the dynamic braking system is designed to provide the full-load retarding torque (worst case), then in lighter load conditions sufficient braking effort will be provided.

Torque is directly related to the armature current by the field current and the constant (K) as shown earlier (2). For separately excited systems, the field current is constant. The proportionality constant, K, can be calculated using nameplate information and empirical data. The required braking torque is simply the full-load imbalanced weight applied at the drum radius through the drive transmission ratio. This steady-state dynamic braking torque can be plotted as a vertical line (Idb) on the terminal characteristic curves, (figure 8) since it is a function of the armature current.

The other design consideration is the desired dynamic braking speed. Dynamic braking does not stop the cage but only limits the final speed to a designed value. Since the dynamic braking system is operating under the condition of mechanical brake failure, the overhauling load will not come to rest until it reaches the bottom of the shaft. Elevators have buffers installed at the bottom of the shaft to reduce the impact of the overhauling load. These buffers are designed to decelerate a cage traveling at 115 percent of rated speed at an average retardation rate of not more than 32.2 ft/s2 (1 g force).[4] Deceleration rates less then 16.1 ft/s2 (1 g force) are internationally accepted to be safe. Therefore the designed dynamic braking speed should be no more than 50 percent of rated speed to limit the deceleration rate to less than 1/2 g force when the cage strikes the buffers.

Once the final braking speed is determined (e.g. 50 percent), the corresponding speed curve is selected from figure 8. A load line (Rdb) can be drawn through the intersection of the torque and speed curve. This resistance load line (Rdb) defines the possible operating points of the dynamic braking system. The initial dynamic braking effort produced when the hoisting system is traveling at rated speed is about twice as large as the steady-state retarding torque as shown by the initial dynamic braking torque line (Io) in figure 8. This initial braking torque rapidly decelerates the hoisting system along the load line (Rdb), down to the steady-state operating point.

Slower final braking speeds can be achieved by decreasing the dynamic braking (DB) resistor to provide a greater margin of safety. However, systems designed to provide the required retarding torque at very low speeds will inherently provide excessive compound braking torques at rated speed. Lowering the DB resistance to provide the required steady-state retarding torque at 25 percent of rated speed will deliver over three times greater braking torque when the cage is traveling at rated speed. This braking torque, when combined with the mechanical braking force, may produce an excessive deceleration rate (i.e. greater than 1/2 g force).

In general, a resistor with an electrical power rating equal to the mechanical horsepower rating of the drive motor will provide sufficient power dissipation for dynamic braking applications. A power rating calculated in this manner will typically incorporate a safety factor.

The separately excited dynamic braking system requires an external power source. If a mine power system interruption or drive power system failure occurs, the separately excited dynamic braking system will not provide retarding torque.

Self-Excited Dynamic Braking System

A self-excited dynamic braking system will generate retarding torque without an external power source. Therefore, in the event of mechanical brake and power system failure, the dynamic braking system can provide sufficient retarding effort to safely lower the overhauling load. The connections for a self-excited dynamic braking system are shown in figure 9.

Figure 9. Self-Excited Dynamic Braking System

Up to this point, the effect of hysteresis and the magnetic retention of the field poles has been neglected to simplify the analysis. However, the operation of a self-excited dynamic braking system is dependent on the residual magnetism that remains in the field pole iron after the field winding is de-energized. The magnetization curves illustrating the effect of this magnetic retention are shown in figure 10.

Figure 10. Self-Excited Magnetization Curves, Under Load,
for Various Percentages of Rated Speed

At rated speed or slower, the generated voltage is directly proportional to the speed as shown for (1). Under load and overspeed conditions, the generated voltage may be less than proportional to the speed due to armature reaction. This effect is shown graphically as a dropping of the generated voltage, eg, during maximum armature current (near the knee of the saturation region) for the 125-percent speed magnetization curve. The saturating effect of the armature reaction is a factor which limits the braking current (torque) during voltage build-up under transient conditions.

The build-up of voltage occurs during the following situation. If the hoisting machine is stopped and no power is available when the mechanical brakes fail, the dynamic braking system must self-excite to generate braking torque. When the brakes fail, the hoisting machine is accelerated by the weight of the overhauling load. The acceleration rate is initially limited by the inertia of the hoisting system and frictional forces. As the armature rotates through the residual magnetic field of the permanently magnetized field poles, a small voltage is generated. This is shown graphically by the voltage intercept of the magnetization curve in figure 10. This small voltage is imposed across the field winding, which creates a small field current. The braking system should be designed so that the flux produced by the resulting ampere-turns of the field winding add to the residual flux. This positive feedback will produce progressively greater voltages, field currents, and retarding torques. The final value of armature voltage and field current are governed by the field, armature, and dynamic-braking load resistances of the circuit and the driving mechanical torque.

If the field-resistance line is below and parallel to the linear portion of the magnetization curve, as shown for the 50-percent rated speed curve, then a very underdamped system response is obtained. The system will oscillate around the required torque and final speed. A stable operating point is obtained where the field resistance line intersects the magnetization curve well into the saturated region.

The steady-state operating point is governed by the following equations:

Vdb      =      eg - Ra ia     V (7)
Vdb      =      Rf if     V (8)

The field and armature voltage drops in (7) and (8) that define the steady-state operating point are shown in figure 10 for the 100-percent rated speed curve. The steady-state dynamic braking voltage (Vdb) for a given armature current ia has a value such that the field current If (8) produces a generated voltage eg (from the magnetization curve), which is Ra ia greater than Vdb (7). The armature current ia, which corresponds to any value of dynamic braking voltage Vdb, can be found by dividing the vertical distance between the magnetization and field-resistance curves in figure 10 by Ra. The dynamic braking load current is:

idb      =      ia - if      A (9)

Figure 11 shows the typical terminal characteristic for a self-excited dynamic braking system. Note that the maximum steady-state dynamic braking current is limited by the maximum vertical separation between the magnetization and field-resistance curves (figure 10). The saturation effect of armature reaction will limit the maximum dynamic braking current as shown for the 125-percent rated speed curve in figure 11.

Figure 11. Terminal Characteristics of a Self-Excited
Generator for Various Percentages of Rated Speed

The dynamic braking voltage (Vdb) intercept corresponds to no-load terminal voltage (Rdb = omega OMEGA, no retarding torque). The dynamic braking current, idb, intercept occurs when the armature terminals are shorted (Rdb = 0 OMEGA; residual retarding torque). The maximum retarding torque occurs for the operating point at the peak of the dynamic braking current. As discussed earlier, optimum dynamic retarding effort is not always desirable when the effects of compound braking are considered.

The retarding torque is directly proportional to the field flux as discussed earlier. Self-excited dynamic braking systems are usually operated in the saturated region of the magnetization curve (figure 10). This corresponds to the upper (linear) portion of the terminal characteristic curve in figure 11. In this highly saturated region the field flux is relatively constant over a wide range of load currents and speeds. Therefore the field flux can be assumed to be constant for the purpose of calculating retarding torque from (2).

The retarding torque is also proportional to the armature current. The armature current is the sum of the field and load currents (9). The field current can be neglected, with minimal error, if it is small with respect to the load current. Therefore the armature current, ia, would equal the load current, idb. The assumption of constant field flux and negligible field current with respect to the load current allows the required steady-state retarding torque line (Idb) to be drawn directly on the load-current axis as shown earlier for the separately excited system.

The desired speed curve is selected (typically 50 percent), and a load line is drawn through the intersection of the required torque and speed curve. The slope of the load line is the approximate ohmic value of the dynamic braking resistor.

Error is introduced into the calculation of the resistor load line by the simplifying assumptions. The assumption of negligible field current with respect to the load current may produce a significant error for systems operating in magnetic saturation. The error is directly proportional to the ratio of the load resistance to field resistance, Rdb /Rf.

The armature current is the sum of the field current if and load current idb (9). Neglecting the field current drain from the armature circuit will result in a larger retarding torques than indicated by the vertical torque line (Idb) drawn on the terminal characteristic curve (figure 11). Thus the final dynamic braking speed will be slower than calculated for the dynamic braking resistance. Therefore the predicted operating point for self-excited dynamic braking system should only be used as a general design guideline.

Case Study: Dynamic Braking Installation and Testing

Performance testing of dynamic braking has been conducted on several mine hoisting systems. A case study of a service elevator that had dynamic braking installed and was subsequently tested will be presented for illustrative purposes.


An elevator accident occurred at a western Pennsylvania coal mine due to a mechanical brake failure. The counterweight fell to the bottom of the 400-ft shaft, causing the cage to overspeed and crash into the headframe. The cage was unoccupied at the time of the accident. The elevator was out of service for several months due to the severity of the damage.

The governor overspeed tripped and attempted to set the safety catches. However, the wedge design of the safety catches only provide braking effort when the cage is traveling down. The installation of dynamic braking was recommended to prevent this type of accident in the future.

Dynamic Braking Installation

A separately or self-excited type of dynamic braking system was installed on the elevator. The equipment needed for the modification included a three-pole loop contactor with two auxiliary contacts (2 NO, 1 NC, 2 NC aux.), a dynamic braking resistor, a single-phase rectifier bridge, and a drive fault relay with six contacts (2 NO, 4 NC). A schematic diagram of the dynamic braking control circuit is shown in figure 12.

Figure 12. Dynamic Braking Control Schematic

The elevator was a gearless drive roped 2:1. It operated at 600 ft/min with a rated capacity of 9000 pounds and was 40 percent over-counterweighted. The nameplate ratings of the elevator were: 115 hp, 127 r/min (600 ft/min cage speed), armature; 407 V, 234 A, field; 17.1 A, 8.61 OMEGA at 25°C. The motor armature was powered by a three-phase full-wave reversing SCR converter. The motor field current was supplied by an SCR-controlled single-phase half-wave rectifier.

When the mechanical brakes were called to set, the M contactor dropped out and disconnected the armature from the power supply, and also applied the dynamic braking resistor across the motor armature. When the field power supply was operative, the drive-OK (DROK) relay was picked-up and the field was supplied with normal standing field current (10 A).

When a total power loss occurred, the dynamic braking resistor was connected across the armature, the DROK relay dropped out, and regenerative braking current was supplied to the motor field. The rectifier bridge in the regenerative field power supply insures the generated amp-turns add to the residual magnetic field for either direction of cage travel. The normally closed DROK contacts were installed to isolate the bridge rectifier from the motor armature and field power supplies.

Dynamic Braking Tests

The dynamic braking tests were designed to demonstrate the response of the hoisting system to various forcing functions. The dynamic braking system was tested under the following conditions:

1) brake failure, power supply operative, starting at rest,
2) brake failure, power supply operative, running at full speed,
3) brake failure, power failure, starting at rest
4) brake failure, power failure, running at full speed,
5) compound braking, running full speed.

These five test conditions were performed with an empty cage (no load) and then repeated with a fully loaded cage (9000 lb).

The armature current, armature voltage, field current, and motor speed were recorded on a strip chart during the tests. The polarity of the signals corresponds to the four quadrants of operation defined in figure 1.

Test Condition 1: A sample strip chart recording for test condition 1, under full load (9000 lb) with a 1.60 - OMEGA dynamic braking resistance, is shown in figure 13. The 9000 lb of load on the cage, with 40-percent over-counterweighting, provided 5400 lb of downward (cage) accelerating force. The speed signal shows the cage accelerated to 330 ft/min before the dynamic braking torque stabilized at the final speed of 280 ft/min (47 percent of rated speed). When the cage was moving, as detected by the tachometer signal, about 5 s after the brakes were called to set, the control logic interrupted the control power (DROK dropped out). Therefore the dynamic braking configuration switched from separately excited to self-excited. In the self-excited mode, and fully loaded, the field was driven into saturation (greater than 17.1 A) as shown by the field current signal of figure 13.

Figure 13. Test Condition (Brake Failure, Power Supply
Operative, Starting at Rest): 9000-lb Cage Load, Rdb = 1.60 OMEGA

Test Condition 2: A sample strip chart recording for test condition 2-no load with a 0.4 - OMEGA dynamic braking resistance-is shown in figure 14. Since the motor was running at rated speed, armature and field currents were established to produce immediate deceleration from rated speed. When dynamic braking was initiated, a large transient armature current (650-A peak) was generated by the relatively low dynamic braking resistance of 0.4 OMEGA. This large armature current generated excessive retarding torque and caused the wire ropes to break traction with the drive sheave. The drive sheave decelerated at an initial rate of 18 ft/s2. When the ropes regained traction with the drive sheave, the motor accelerated briefly before slowing down to the steady-state speed of 150 ft/min.

Figure 14. Test Condition 2 (Brake Failure, Power Supply
Operative, Running at Full Speed): No Cage Load, Rdb = 0.4 OMEGA

Test Condition 3: A sample strip chart recording for test condition 3-no load with 1.60 OMEGA of dynamic braking resistance-is shown in figure 15. An empty cage with 40-percent over-counterweighting provided 3600 lb of upward cage-accelerating force. When the mechanical brakes were defeated, the cage accelerated upward at an initial rate of about 1.5 ft/s2. As the cage accelerated, the armature windings rotated rapidly through the weak residual magnetic field of the permanently magnetized field poles. Thus a small amount of armature current was generated, which divided between the dynamic braking resistor and the field winding. The field current increased the strength of the magnetic field, which in turn increased the generated armature current. This positive feedback continued, causing the field current to build on the generated armature current until a sufficient retarding torque was developed at 700 ft/min and the cage began to decelerate. The field and armature currents then began to decrease as the cage decelerated. The cage slowed down to a steady-state speed of 200 ft/min with an underdamped response.

Figure 15. Test Condition 3 (Brake Failure, Power Failure,
Starting at Rest): No Cage Load, Rdb = 1.60 OMEGA

The peak speed reached during the self-excitation process was primarily a function of the time constant of the inductive motor field winding and the acceleration rate of the cage. The inductance of the field winding was fixed; however, the acceleration rate of the cage was a function of the load inertia and the imbalance between the cage and counterweight. Larger imbalance conditions would produce greater acceleration rates. Faster acceleration rates may produce excessive maximum speed situations during the self-excitation process. The maximum acceleration rate for personnel load conditions occurs when one person is being hoisted. As more persons are added to the cage (up to the rated personnel capacity) the load imbalance between the cage and counterweight is reduced, thereby reducing the acceleration rate.

When test condition 3 was performed, with the maximum material load of 9000 lb, the cage accelerated at 2.5 ft/s2. The rapid acceleration rate caused the cage to reach an excessive speed (870 ft/min) and trip the safety catches before the dynamic braking effort had time to develop and decelerate the cage.

The dynamic braking system would self-excite more rapidly with the dynamic braking resistor disconnected. The dynamic braking resistor could be applied across the armature, to generate retarding torque, once the armature and field currents were generated. However, this approach would complicate the control scheme and reduce the simplicity and reliability of the dynamic braking system.

It should be noted, however, that in given time the dynamic braking effort would be developed for the fully loaded situation. In addition, the cage was equipped with safety catches that operate and stop the cage when it is falling.

Test Condition 4: A sample strip chart recording for test condition 4-full load with 1.60 OMEGAof dynamic braking resistance-is shown in figure 16.

Figure 16. Test Condition 4 (Brake Failure, Power Failure,
Running Full Speed): 9000-lb Cage Load, Rdb = 1.60 OMEGA

The main disconnect switch was opened while the elevator was running at 600 ft/min with 9000 lb of load on the cage. Dynamic braking effort was available the instant the power was interrupted, since armature and field currents were flowing in the motor. The cage decelerated from 600 ft/min to a final speed of 300 ft/min (50 percent of rated speed) with no oscillations.

This was the worst-case power-loss situation since the drive motor was regenerating maximum steady-state power into the ac power line to dissipate the energy of the falling fully loaded cage. When the ac power line was opened, no path was available to conduct the regenerated power; therefore a large transient voltage was generated. Since the energy could not be dissipated in the power system impedance, the transient energy coupled through the control transformer and blew two SCR drive control fuses and one fuse on the primary side of the control transformer. The fully loaded cage hit the buffers at the final dynamic braking speed of 300 ft/min without sustaining any damage. The three fuses were replaced and the elevator operated properly.

Test Condition 5: Excessive compound braking, running at full speed, occurred when the dynamic braking effort was optimized by installing a dynamic braking resistance of 0.4 OMEGA.

The dynamic braking force acting alone produced a drive sheave decelerating rate of 18 ft/s2 (figure 14). When this braking force was added to the mechanical braking force, the elevator drive decelerated at 26 ft/s2. The dynamic braking effort was reduced by increasing the dynamic braking resistance from 0.4 to 1.60 OMEGA. This increased the slope of the load line (Rdb) on the terminal characteristic curve (figures 8 and 11). Thus armature current was reduced, thereby reducing the retarding torque. The resulting deceleration rates, during compound braking, for lowering and hoisting the empty cage were 7.1 and 10.4 ft/s2, respectively. This was only a 1 ft/s2 increase over the deceleration rate of the mechanical braking system acting alone. The steady-state speed increased from 150 to 200 ft/min.


Several dynamic braking control configurations have been discussed. Other configurations could be designed; however, their increased performance is offset by the additional complexity of installation, greater cost, and reduced reliability. A dynamic braking system can be designed to safely lower an overhauling load, even under simultaneous multiple failures of the mechanical braking system and power system. These simple dynamic-braking control configurations provide an increased margin of safety at a modest expense.



W. J. Helfrich, "Island Creek Coal Company V. P. - 5 Mine, Mine Safety and Health Administration, Mine Electrical Systems Division Investigative Report No. C080978, August, 1978.

Safe Manriding in Mines, First Report of the National Committee for Safety of Manriding in Shafts and Unwalkable Outlets, Her Majesty's Stationery Office, Health and Safety Executive, London, UK, 1976.

Application Manual DC Elevator Equipment, Erie, Pennsylvania: General Electrical Company.

Safety Code for Elevators and Escalators, New York: ANSI/ASME A17.1-1984, Rule 201.4, pg. 47.

I. L. Kosow, Electric Machinery and Transformers, New Jersey: Prentice-Hall, 1972, ch. 3, pp. 91-92.