Advance Measurement Techniques in Turbomachines

*Fangyuan Lou*

## **Abstract**

This chapter focuses on advanced measurement techniques that have been used in applications of turbomachines including temperature measurements, pressure measurements, velocity measurements, and strain/stress measurements. Though the measurement techniques are fundamentally the same as those used in other applications, the unique features associated with turbomachines place challenges in implementing these techniques. This chapter covers the fundamental working principles of individual measurement technique as well as the highlights of its application in turbomachines.

**Keywords:** measurement techniques, temperature-sensitive paint, pressure-sensitive paint, laser Doppler velocimetry, particle image velocimetry, hot-wire anemometry, strain gauges, nonintrusive stress measurement systems

## **1. Introduction**

Turbomachine consists a wide and diverse class of devices that have been used in air, land, and sea. Below is a list of representative applications for turbomachines:


The flow in turbomachines is highly three-dimensional, turbulent, and inherently unsteady. The unsteady nature of the flow in turbomachines is a result of work exchange between the machine and its working fluid. These complex flow phenomena affect the performance and operability. The interactions between the flow and hardware structures can result in undesired noise, vibration, and sometimes failure of the machine. On one hand, enhanced understanding of the complex flow

phenomena in turbomachines is essential for the development of better turbomachines in the future and, thus, requires experimental benchmark data associated with the flow including velocity, pressure, and temperature. On the other hand, better monitoring of the health status of the rotating groups (i.e., stress of rotors) is also of great importance for failure prevention.

Extensive experimental studies have been performed during the past few decades to investigate the complex flow as well as fluid-structure interactions in turbomachines. This chapter discusses the advanced techniques as well as highlights the challenges in implementing these techniques.

### **2. Velocity and turbulence measurements**

This section discusses three techniques for velocity and turbulence measurements in applications of turbomachines including hot-wire anemometry, laser Doppler velocimetry (LDV), and particle image velocimetry (PIV). A brief introduction of the working principles, features associated with the measurement technique, and the challenges for implementation in turbomachines are presented. Lists of previous studies in the open literature applying these measurement techniques to turbomachines are also provided.

#### **2.1 Thermal anemometer**

Hot-wire anemometry is an intrusive measurement technique that provides instantaneous velocity and turbulence measurements. It allows characterizing high-frequency flow structures at relatively inexpensive cost when compared with alternative approaches such as laser Doppler velocimetry and other optical techniques. Below highlights the features of the hot-wire anemometry.

*Intrusive velocity measurement.* In contrast to techniques for the measurement of flow velocities employing probes such as pressure tubes or hot wires, the LDV technique features a nonintrusive nature eliminating the disturbances introduced by the presence of the probes.

*Direct velocity measurement*. The hot-wire anemometry does not require particle tracers and provides direct measurements of the fluid velocity and turbulence.

*Point measurement*. Hot-wire anemometry is a point-based measurement, and the measurement volume is determined by the dimensions of the employed wires.

A hot-wire probe consists of a short-length (on the magnitude of millimeter), fine-diameter (on the magnitude of micrometer) wire that is attached to two prongs. The technique relies on the changes in heat transfer between the heated wire and the fluid the wire is exposed. Heat is introduced in the sensor by Joule heating and is lost primarily by forced convection. A significant parameter that controls the operation of the sensor is the relative difference in temperature between the heated up wire and the flow, which is related to the overheat ratio of the sensor. Changes in the flow properties, such as velocity, density, temperature, and humidity, will cause a corresponding change in the heat transfer from the wire which can be detected and measured. Hot wires are typically run in either a constant temperature (CT) mode or constant current (CC) mode. A sketch of circuit for thermal anemometry operating in constant temperature mode is shown in **Figure 1**. Constant temperature anemometry utilizes a rapid-response servo circuit coupled with the Wheatstone bridge amplifier to control the applied voltage and maintain a constant wire resistance, which in turn maintains a constant wire temperature. It eliminates the effect of thermal inertial of the wire, as well as the system response time, and, thus, provides a better frequency response compared to the constant current

**27**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

offered by the hot-wire anemometry.

**Figure 1.**

lence and Reynolds stress measurements.

studies in applications of turbomachines.

**2.2 Laser Doppler velocimetry (LDV)**

surement technique:

by the presence of the probes.

tracer particle being added to the flow.

operation. Most velocity and turbulence measurements are acquired in this manner. Attributed to the extremely small thermal mass of the wire, this technique allows detection of very-high-frequency fluctuations in the flow. This is another advantage

*Circuit diagram of a thermal anemometer operating at constant temperature mode.*

The hot wires can be configured in the manner of single wire, cross wires, and triple wires. The single wire is primarily useful for mean flow quantity measurements and is not as accurate as triple wire simultaneous measurements for turbu-

Because of the small size and high-frequency response offered by hot-wire anemometer, the technique has been used extensively in investigations of the flow fields in turbomachines. While early investigations using hot-wire/hot-film probes were mostly qualitative, substantial amounts of quantitative investigations have also been carried out in the past few decades on various aspects, using many different types of hot-wire and hot-film probes. **Table 1** provides a list of representative

Laser Doppler velocimetry is an optical nonintrusive technique which measures the instantaneous velocity at a given point in a flow field. It was first developed by Yeh and Cummins in 1964 and is now a well-established technique. This technique has been widely used in all kinds of fluid flow applications, including laminar flow, turbulent flow, flow inside turbomachinery, and flow inside combustion chambers. Because of its high accuracy, it is also used as a benchmark validation tool for planar velocimetry techniques (i.e., PIV, PTV). Below lists the features in the LDV mea-

*Nonintrusive velocity measurement*. In contrast to techniques for the measurement of flow velocities employing probes such as pressure tubes or hot wires, the LDV technique features a nonintrusive nature eliminating the disturbances introduced

*Indirect velocity measurement*. The LDV technique measures the velocity of a fluid element indirectly in most of cases by means of the measurement of the velocity of

*Point measurement*. The same as hot wire, the LDV is a point-based measurement

A sketch of the LDV working principle is shown in **Figure 2**. The system shown in the sketch is a 1-D LDV system operated in backscattering mode, but the working theory is the same for all the LDV systems. A single laser beam is emitted from a

technique that characterizes the velocity in its measurement volume.

**Figure 1.**

*Rotating Machinery*

also of great importance for failure prevention.

the challenges in implementing these techniques.

**2. Velocity and turbulence measurements**

niques to turbomachines are also provided.

**2.1 Thermal anemometer**

by the presence of the probes.

phenomena in turbomachines is essential for the development of better turbomachines in the future and, thus, requires experimental benchmark data associated with the flow including velocity, pressure, and temperature. On the other hand, better monitoring of the health status of the rotating groups (i.e., stress of rotors) is

Extensive experimental studies have been performed during the past few decades to investigate the complex flow as well as fluid-structure interactions in turbomachines. This chapter discusses the advanced techniques as well as highlights

This section discusses three techniques for velocity and turbulence measurements in applications of turbomachines including hot-wire anemometry, laser Doppler velocimetry (LDV), and particle image velocimetry (PIV). A brief introduction of the working principles, features associated with the measurement technique, and the challenges for implementation in turbomachines are presented. Lists of previous studies in the open literature applying these measurement tech-

Hot-wire anemometry is an intrusive measurement technique that provides instantaneous velocity and turbulence measurements. It allows characterizing high-frequency flow structures at relatively inexpensive cost when compared with alternative approaches such as laser Doppler velocimetry and other optical tech-

*Intrusive velocity measurement.* In contrast to techniques for the measurement of flow velocities employing probes such as pressure tubes or hot wires, the LDV technique features a nonintrusive nature eliminating the disturbances introduced

*Direct velocity measurement*. The hot-wire anemometry does not require particle

tracers and provides direct measurements of the fluid velocity and turbulence. *Point measurement*. Hot-wire anemometry is a point-based measurement, and the measurement volume is determined by the dimensions of the employed wires. A hot-wire probe consists of a short-length (on the magnitude of millimeter),

fine-diameter (on the magnitude of micrometer) wire that is attached to two prongs. The technique relies on the changes in heat transfer between the heated wire and the fluid the wire is exposed. Heat is introduced in the sensor by Joule heating and is lost primarily by forced convection. A significant parameter that controls the operation of the sensor is the relative difference in temperature between the heated up wire and the flow, which is related to the overheat ratio of the sensor. Changes in the flow properties, such as velocity, density, temperature, and humidity, will cause a corresponding change in the heat transfer from the wire which can be detected and measured. Hot wires are typically run in either a constant temperature (CT) mode or constant current (CC) mode. A sketch of circuit for thermal anemometry operating in constant temperature mode is shown in **Figure 1**. Constant temperature anemometry utilizes a rapid-response servo circuit coupled with the Wheatstone bridge amplifier to control the applied voltage and maintain a constant wire resistance, which in turn maintains a constant wire temperature. It eliminates the effect of thermal inertial of the wire, as well as the system response time, and, thus, provides a better frequency response compared to the constant current

niques. Below highlights the features of the hot-wire anemometry.

**26**

*Circuit diagram of a thermal anemometer operating at constant temperature mode.*

operation. Most velocity and turbulence measurements are acquired in this manner. Attributed to the extremely small thermal mass of the wire, this technique allows detection of very-high-frequency fluctuations in the flow. This is another advantage offered by the hot-wire anemometry.

The hot wires can be configured in the manner of single wire, cross wires, and triple wires. The single wire is primarily useful for mean flow quantity measurements and is not as accurate as triple wire simultaneous measurements for turbulence and Reynolds stress measurements.

Because of the small size and high-frequency response offered by hot-wire anemometer, the technique has been used extensively in investigations of the flow fields in turbomachines. While early investigations using hot-wire/hot-film probes were mostly qualitative, substantial amounts of quantitative investigations have also been carried out in the past few decades on various aspects, using many different types of hot-wire and hot-film probes. **Table 1** provides a list of representative studies in applications of turbomachines.

### **2.2 Laser Doppler velocimetry (LDV)**

Laser Doppler velocimetry is an optical nonintrusive technique which measures the instantaneous velocity at a given point in a flow field. It was first developed by Yeh and Cummins in 1964 and is now a well-established technique. This technique has been widely used in all kinds of fluid flow applications, including laminar flow, turbulent flow, flow inside turbomachinery, and flow inside combustion chambers. Because of its high accuracy, it is also used as a benchmark validation tool for planar velocimetry techniques (i.e., PIV, PTV). Below lists the features in the LDV measurement technique:

*Nonintrusive velocity measurement*. In contrast to techniques for the measurement of flow velocities employing probes such as pressure tubes or hot wires, the LDV technique features a nonintrusive nature eliminating the disturbances introduced by the presence of the probes.

*Indirect velocity measurement*. The LDV technique measures the velocity of a fluid element indirectly in most of cases by means of the measurement of the velocity of tracer particle being added to the flow.

*Point measurement*. The same as hot wire, the LDV is a point-based measurement technique that characterizes the velocity in its measurement volume.

A sketch of the LDV working principle is shown in **Figure 2**. The system shown in the sketch is a 1-D LDV system operated in backscattering mode, but the working theory is the same for all the LDV systems. A single laser beam is emitted from a


#### **Table 1.**

*Representative studies that have used thermal anemometer technique in studying the flow field of turbomachines.*

**29**

**Table 2.**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

**Author(s) Year Type Type of machine Subject of study**

Murthy and Lakshminarayana [19] 1986 1-D Axial compressor Rotor tip region flow Beaudoin et al. [20] 1992 2-D Centrifugal pump Effects of orbiting impeller

Farrell and Billet [22] 1994 na Axial pump Tip vortex cavitation

Faure et al. [28] 2001 2-D Axial compressor Flow structure Van Zante et al. [29] 2002 2-D Axial compressor Blade row interactions

Ibaraki et al. [30] 2003 2-D Centrifugal impeller Impeller flow Higashimori et al. [31] 2004 2-D Centrifugal impeller Impeller flow field

Ibaraki et al. [34] 2009 2-D Centrifugal impeller Impeller flow field

*Representative studies that have used LDV technique in studying the flow field of turbomachines.*

Abramian and Howard [24] 1994 1-D Centrifugal impeller Impeller relative flow field Zaccaria and Lakshminarayana [25] 1997 2-D Axial turbine Rotor passage flow field Adler and Benyamin [26] 1999 2-D Axial turbine Stator wake propagation Ristic et al. [27] 1999 3-D Axial turbine 3-D flow field downstream of

rotor

compressor

compressor

Pierzga and Wood [17] 1985 1-D Axial fan rotor 3-D flow field in a transonic rotor Strazisar [18] 1985 1-D Axial fan rotor Flow structure in transonic fan

Rotor passage relative flow

rotor

3-D flow structure

Flow structure

rotor turbine

Impeller discharge flow

Diffuser flow

Wisler and Mossey [16] 1973 1-D Axial compressor

Hathaway et al. [21] 1993 3-D Centrifugal

Fagan and Fleeter [23] 1994 1-D Centrifugal

Schleer et al. [33] 2004 2-D Centrifugal

Gooding et al. [35] 2019 3-D Centrifugal

laser head operating in continuous mode and then enters into the optical transmitter. Inside the transmitter, this single beam is split and frequency shifted using the beam splitter (BS), an achromatic lens, and a Bragg cell. Pairs of monochrome laser beams (depending on the number of velocity components needs to be measured: one pair for a 1-D system, two pairs for a 2-D system, and three pairs for a 3-D system) generated by the transmitter are then conveyed to the optical probes using fiber cables. The laser beams coming out of the probe intersect at the focal point of the front lens. At this focal point, at which the measurement volume is located, an ellipsoidal volume with bright and dark fringe patterns is formed by the interference of the laser beams. As flow particles traverse through this measurement volume, the backscattered light is collected by the receiving optics inside the probe and further processed by a burst spectrum analyzer. Inside the spectrum analyzer, the time intervals for the burst traveling through the bright and dark patterns are measured. Those measured time intervals, combined with the known distance between the adjacent bright and dark strips, yield the calculation of velocity. This nonintrusive feature of the technique attracted the attention of experimentalist in the field of turbomachines soon after it was introduced in the 1960s. In addition to being nonintrusive, it allowed the velocity measurements in the rotating

Faure et al. [32] 2004 3-D Axial compressor Reynolds stresses measurements

compressor

compressor

**Figure 2.** *Sketch of 1-D backscattering LDV system.*

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*


#### **Table 2.**

*Rotating Machinery*

Lakshminarayana and

Lakshminarayana [3]

Poncet [1]

Hah and

Ristic and

**Table 1.**

*turbomachines.*

Lakshminarayana [9]

Gorton and Lakshminarayana [2]

**Author(s) Year Sensor type Type of machine Subject of study**

Axial inducer Rotor wakes

rotor wake

Axial compressor Turbulence intensity and length scale

Axial compressor 3-D wake decay and

on rotor wake

secondary flows

Diffuser flow

during spike-type stalling

spike-type stalling

Unsteady boundary layer

Rotation and blade incidence

1976 Triple wire Axial inducer Mean flow and turbulence

1980 Triple wire Axial compressor Freestream turbulence on a

Axial compressor and turbine

1998 Cross wire Axial turbine 3-D boundary layer

compressor

turbine

1974 Singe and cross wire

> film, single sensor

mounted hot film/hot wire

Hsu and Wo [8] 1997 Slanted hot wire Axial compressor Unsteady wake

Furukawa et al. [10] 1998 Hot wire Diagonal flow rotor Tip flow field

Sentker and Riess [11] 2000 Split hot film Axial compressor Wake-blade interaction Velarde-Suarez et al. [12] 2001 Cross wire Centrifugal fan Unsteady flow

Goodhand and Miller [14] 2011 Single wire Axial compressor Boundary layer development

Weichert and Day [15] 2014 Single wire Axial compressor Tip region flow during

*Representative studies that have used thermal anemometer technique in studying the flow field of* 

wire)

Hodson et al. [4] 1994 Hot film Low-pressure

Pinarbasi [13] 2008 Triple wire Centrifugal

Camp and Shin [5] 1995 Hot wire, hot

Witkowski et al. [6] 1996 Hot film (triple

Halstead et al. [7] 1997 Surface-

**28**

**Figure 2.**

*Sketch of 1-D backscattering LDV system.*

*Representative studies that have used LDV technique in studying the flow field of turbomachines.*

laser head operating in continuous mode and then enters into the optical transmitter. Inside the transmitter, this single beam is split and frequency shifted using the beam splitter (BS), an achromatic lens, and a Bragg cell. Pairs of monochrome laser beams (depending on the number of velocity components needs to be measured: one pair for a 1-D system, two pairs for a 2-D system, and three pairs for a 3-D system) generated by the transmitter are then conveyed to the optical probes using fiber cables. The laser beams coming out of the probe intersect at the focal point of the front lens. At this focal point, at which the measurement volume is located, an ellipsoidal volume with bright and dark fringe patterns is formed by the interference of the laser beams. As flow particles traverse through this measurement volume, the backscattered light is collected by the receiving optics inside the probe and further processed by a burst spectrum analyzer. Inside the spectrum analyzer, the time intervals for the burst traveling through the bright and dark patterns are measured. Those measured time intervals, combined with the known distance between the adjacent bright and dark strips, yield the calculation of velocity.

This nonintrusive feature of the technique attracted the attention of experimentalist in the field of turbomachines soon after it was introduced in the 1960s. In addition to being nonintrusive, it allowed the velocity measurements in the rotating reference frame without having to use complex rotating probe traverse or data transmission mechanisms (i.e., slip ring or telemetry system). A list of representative studies using LDV for measurements in turbomachines is provided in **Table 2**.

One of the earliest applications of LDV to turbomachinery was conducted by Wisler and Mossey [16] to measure the relative velocity across the first-stage rotor blade row using a single-component LDV system. The flow was seeded by spray atomizing a dilute water suspension of 1-μm-diameter polystyrene latex particles. A sketch of the experimental setup and a sample contour plot of relative velocity within the rotor passage at mid-span (50%) are shown in **Figure 3**. In addition, a sketch of experimental setup for three-component LDV in a centrifugal compressor is presented in **Figure 4**. As summarized in the table, majority of the investigations involving LDV have been performed in a stationary frame of reference. To measure the flow field in rotors, the rotor passage period has been discretized into bins, each with a finite time interval. The results in each bin were ensemble averaged to obtain the mean velocity and turbulence parameters across the rotor passage. To reach convergence in mean velocity and turbulence parameters, a large data set per bin is favored which requires a larger bin size. However, an increase in the bin size introduces the effects of spatial variations in the flow structure. An alternative approach is to conduct measurements in the rotating frame of reference. However, this makes the experimental very challenging, and the only study reported in the open literature of this category was performed by Abramian and Howard [24]. The experiment was conducted in a centrifugal impeller using a Dove prism to transfer the laser beams to the rotating frames of reference.

There are challenges in implementing LDV to turbomachines. Generally speaking, the challenges can be categorized into the optical accessibility-related issues and particle related. Typically, the three-dimensional twisted rotor blades make it difficult to shine laser beams to the interested measurement locations and require the LDV system operating in backscatter mode. Comparing to the favorable forward-scatter configuration, the signal-to-noise ratio of backscatter mode is commonly one to three orders of magnitude smaller. Additionally, the signal-to-noise ratio gets further deteriorated at measurement locations close to metal surfaces due to reflections and in applications of curved optical windows due to the distortion of laser beams through the windows. These distortions increase the uncertainty of the measurements by deforming the measurement volume and changing the measurement location. It is also challenging to deliver particles to target measurement locations due to the strong secondary flow in turbomachines.

## **2.3 Particle image velocimetry (PIV)**

In addition to LDV, particle image velocimetry is another *nonintrusive* technique for velocity measurements. The same as LDV, PIV is also an *indirect* measurement technique and requires tracer particles. Different from LDV and thermal anemometer which are point-based measurement technique, PIV offers *full-field* measurements and allows mapping of large parts of flow field. The working principle of PIV is schematically described in **Figure 5**. The principle of PIV is based on the measurement of the displacement of small tracer particles during a short time interval. This indirect measurement nature requires the tracer particles to be sufficiently small to precisely follow the motion of fluid. The tracer particles are typically illuminated using a thin light sheet generated from pulsed laser head. A pair of images for the illuminated flow field is taken by a digital imaging device, typically a CCD camera. Depending on the number and configuration of camera employed, either 2-D or 3-D flow field could be obtained using cross-correlation analysis to measure the displacement

**31**

**Figure 3.**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

of particles in each small interrogation areas. A single-camera system allows characterization of the two velocity components within the measurement plane, while stereo imaging using two inclined cameras provides all three components

*The LDV setup (a) and measured relative velocity contours within the rotor passage of a low-speed research* 

Effort of implementing PIV in investigations of turbomachinery flow filed has been entertained since the emergence of the technique. Previous researchers have performed both two-dimensional and stereoscopic PIV measurements within various axial and centrifugal turbomachinery facilities. A few highlights of selected previous research are presented, and a more extended set of references is provided in **Table 3**.

of the velocity in the illuminated plane.

*compressor (b) Wisler and Mossey [16].*

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

*Rotating Machinery*

reference frame without having to use complex rotating probe traverse or data transmission mechanisms (i.e., slip ring or telemetry system). A list of representative studies using LDV for measurements in turbomachines is provided in **Table 2**. One of the earliest applications of LDV to turbomachinery was conducted by Wisler and Mossey [16] to measure the relative velocity across the first-stage rotor blade row using a single-component LDV system. The flow was seeded by spray atomizing a dilute water suspension of 1-μm-diameter polystyrene latex particles. A sketch of the experimental setup and a sample contour plot of relative velocity within the rotor passage at mid-span (50%) are shown in **Figure 3**. In addition, a sketch of experimental setup for three-component LDV in a centrifugal compressor is presented in **Figure 4**. As summarized in the table, majority of the investigations involving LDV have been performed in a stationary frame of reference. To measure the flow field in rotors, the rotor passage period has been discretized into bins, each with a finite time interval. The results in each bin were ensemble averaged to obtain the mean velocity and turbulence parameters across the rotor passage. To reach convergence in mean velocity and turbulence parameters, a large data set per bin is favored which requires a larger bin size. However, an increase in the bin size introduces the effects of spatial variations in the flow structure. An alternative approach is to conduct measurements in the rotating frame of reference. However, this makes the experimental very challenging, and the only study reported in the open literature of this category was performed by Abramian and Howard [24]. The experiment was conducted in a centrifugal impeller using a Dove prism to transfer

There are challenges in implementing LDV to turbomachines. Generally speaking, the challenges can be categorized into the optical accessibility-related issues and particle related. Typically, the three-dimensional twisted rotor blades make it difficult to shine laser beams to the interested measurement locations and require the LDV system operating in backscatter mode. Comparing to the favorable forward-scatter configuration, the signal-to-noise ratio of backscatter mode is commonly one to three orders of magnitude smaller. Additionally, the signal-to-noise ratio gets further deteriorated at measurement locations close to metal surfaces due to reflections and in applications of curved optical windows due to the distortion of laser beams through the windows. These distortions increase the uncertainty of the measurements by deforming the measurement volume and changing the measurement location. It is also challenging to deliver particles to target measurement

In addition to LDV, particle image velocimetry is another *nonintrusive* technique for velocity measurements. The same as LDV, PIV is also an *indirect* measurement technique and requires tracer particles. Different from LDV and thermal anemometer which are point-based measurement technique, PIV offers *full-field* measurements and allows mapping of large parts of flow field. The working principle of PIV is schematically described in **Figure 5**. The principle of PIV is based on the measurement of the displacement of small tracer particles during a short time interval. This indirect measurement nature requires the tracer particles to be sufficiently small to precisely follow the motion of fluid. The tracer particles are typically illuminated using a thin light sheet generated from pulsed laser head. A pair of images for the illuminated flow field is taken by a digital imaging device, typically a CCD camera. Depending on the number and configuration of camera employed, either 2-D or 3-D flow field could be obtained using cross-correlation analysis to measure the displacement

the laser beams to the rotating frames of reference.

locations due to the strong secondary flow in turbomachines.

**2.3 Particle image velocimetry (PIV)**

**30**

**Figure 3.**

*The LDV setup (a) and measured relative velocity contours within the rotor passage of a low-speed research compressor (b) Wisler and Mossey [16].*

of particles in each small interrogation areas. A single-camera system allows characterization of the two velocity components within the measurement plane, while stereo imaging using two inclined cameras provides all three components of the velocity in the illuminated plane.

Effort of implementing PIV in investigations of turbomachinery flow filed has been entertained since the emergence of the technique. Previous researchers have performed both two-dimensional and stereoscopic PIV measurements within various axial and centrifugal turbomachinery facilities. A few highlights of selected previous research are presented, and a more extended set of references is provided in **Table 3**.

#### *Rotating Machinery*

**Figure 6** presents sample results from the two-dimensional measurements performed in an axial pump at Johns Hopkins University. The distribution of phase-averaged velocity, vorticity, and turbulent kinetic energy at the mid-span of the second stage was characterized [37]. **Figure 7** shows sample data obtained in a high-speed centrifugal compressor operating both at the design point and during surge [43, 46]. As summarized in the table, majority of these studies insert a periscopic optical probe into the flow for light sheet delivery. This results in invasive measurement and also

**33**

**Figure 6.**

**Table 3.**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

**Author(s) Year Type Type of machine Subject of study** Paone et al. [38] 1989 2-D Centrifugal pump Flow structure

Chu et al. [39, 40] 1995 2-D Centrifugal pump Unsteady flow and pressure

Day et al. [41] 1996 2-D Axial turbine Effect of film cooling on flow

Uzol and Camci [45] 2001 2-D Axial turbine cascade Trailing edge coolant ejection Wernet et al. [46] 2001 2-D Centrifugal compressor Diffuser flow during surge Chow et al. [37] 2002 2-D Axial pump Wake-wake interactions

Uzol et al. [47] 2002 2-D Axial pump Unsteady flow and deterministic

Uzol et al. [51] 2003 3-D Axial pump 3-D wake structure and tip vortex

Sanders et al. [48] 2002 2-D Axial compressor Blade row interactions Estevadeordal et al. [49] 2002 2-D Axial compressor Wake-blade interactions

Woisetschlager et al. [50] 2003 2-D Axial turbine cascade Turbine wake

Lee et al. [52] 2004 3-D Marine propeller Propeller wake Wernet et al. [53] 2005 3-D Axial compressor Tip region flow Yu and Liu [54] 2006 3-D Axial compressor Unsteady flow Ibaraki et al. [55] 2007 2-D Centrifugal compressor Unsteady diffuser flow Estevadeordal et al. [56] 2007 3-D Axial compressor Wake-rotor interactions

Voges et al. [57] 2007 2-D Centrifugal compressor Diffuser flow Voges et al. [58] 2012 3-D Axial compressor Tip region flow Guillou et al. [59] 2012 3-D Turbocharger compressor Impeller inlet flow Gancedo et al. [60] 2016 3-D Turbocharger compressor Impeller inlet flow Bhattacharya et al. [61] 2016 3-D Axial compressor Rotor flow field

*Representative studies that have used PIV technique in studying the flow field of turbomachines.*

*Sample PIV data obtained in an axial pump facility at Johns Hopkins University: Phase-averaged velocity field (top left), turbulent kinetic energy (bottom left), and vorticity (right) at mid-span within an entire stage [37].*

Dong et al. [42] 1997 2-D Centrifugal pump Unsteady flow and noise Wernet [43] 2000 2-D Centrifugal compressor Diffuser flow structure Sinha and Katz [44] 2000 2-D Centrifugal pump Diffuser flow field

fluctuations

structure

stresses

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

*Rotating Machinery*

**Figure 4.**

**32**

**Figure 5.**

*Sketch of typical PIV setup [36].*

**Figure 6** presents sample results from the two-dimensional measurements performed in an axial pump at Johns Hopkins University. The distribution of phase-averaged velocity, vorticity, and turbulent kinetic energy at the mid-span of the second stage was characterized [37]. **Figure 7** shows sample data obtained in a high-speed centrifugal compressor operating both at the design point and during surge [43, 46]. As summarized in the table, majority of these studies insert a periscopic optical probe into the flow for light sheet delivery. This results in invasive measurement and also

*Photo of experimental setup for study of flow in a centrifugal compressor using three-component LDV.*


#### **Table 3.**

*Representative studies that have used PIV technique in studying the flow field of turbomachines.*

#### **Figure 6.**

*Sample PIV data obtained in an axial pump facility at Johns Hopkins University: Phase-averaged velocity field (top left), turbulent kinetic energy (bottom left), and vorticity (right) at mid-span within an entire stage [37].*

**Figure 7.**

*Sample PIV measurements in the diffuser passage of a high-speed centrifugal compressor at both the design point (left) [43] and during a surge (right) [46].*

**Figure 8.**

*Experimental setup for PIV measurements performed at Purdue University (left) and sample results (right) of normalized radial velocity at fixed spanwise locations for stereo reconstructed velocity field [61].*

significantly limits the region of flow field that can be imaged. To address these challenges, a new approach was introduced in a recent study performed at Purdue University in a multistage axial compressor [61]. The same window was used for both laser sheet delivery and image recording. By doing so, it eliminates the presence of invasive probe for light sheet delivery. A sketch of the experimental setup and sample results is shown in **Figure 8**. As illustrated in the figure, the PIV measurements were performed in the second-stage rotor passage (rotor 2). To eliminate light reflections from the blade surface and hub, fluorescent dye with sufficiently separated absorption and emission wavelengths was introduced with the seeding fluid, and lens filters blocking wavelengths below 540 nm were used to filter laser reflections. Slices of normalized radial velocity at fixed spanwise positions were presented to illustrate the development of the tip leakage flow across the rotor passage.

### **3. Pressure-sensitive paints**

Conventionally, surface pressures are measured using hundreds of pressure taps or flush-mounted transducers to obtain a reasonable spatial distribution. This makes

**35**

**Table 4.**

**Figure 9.**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

*Schematic of pressure-sensitive paint measurement system [36].*

Liu et al. [64] 1997 High-speed axial

Lepicovsky and Bencic [70] 2002 Supersonic through flow

*Representative studies that have used PSP technique in turbomachines.*

**Author(s) Year Type of machine Subject of study**

Sabroske et al. [63] 1995 Axial compressor Blade pressure distribution

Blade surface pressure

Effect of change operating conditions

conditions

pressures

cooling

compressor

Engler et al. [67] 2000 Axial turbine Shock movement and corner stall Navarra et al. [68] 2001 Axial compressor Blade surface pressure in transonic

Gregory [71] 2004 Centrifugal compressor Effect of inlet distortion on surface

Narzary et al. [73] 2012 Axial turbine Effect of coolant density on turbine film

Navarra [65] 1997 Axial compressor Blade surface pressure Bencic [66] 1998 Axial fan Blade surface pressure

Gregory et al. [69] 2002 Centrifugal compressor Blade surface pressure

fan

Suryanarayanan et al. [72] 2010 Axial turbine Filming cooling

the measurements time-consuming and expensive. Recently, the introduction of pressure-sensitive paint (PSP) provides a new method for surface pressure measurement. Comparing to the conventional approaches by means of pressure taps or transducers which can only provide data at discrete points and are limited by installation locations, the PSP technique is very attractive; hence, it provides high-spatial-resolution pressure measurements without taps or transducers. The PSP technique is based on covering a surface with luminescent coatings. The luminescence of the coating is dependent on surface static pressure. With proper illumination, the surface pressure

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

*Rotating Machinery*

**34**

**Figure 8.**

**Figure 7.**

*point (left) [43] and during a surge (right) [46].*

significantly limits the region of flow field that can be imaged. To address these challenges, a new approach was introduced in a recent study performed at Purdue University in a multistage axial compressor [61]. The same window was used for both laser sheet delivery and image recording. By doing so, it eliminates the presence of invasive probe for light sheet delivery. A sketch of the experimental setup and sample results is shown in **Figure 8**. As illustrated in the figure, the PIV measurements were performed in the second-stage rotor passage (rotor 2). To eliminate light reflections from the blade surface and hub, fluorescent dye with sufficiently separated absorption and emission wavelengths was introduced with the seeding fluid, and lens filters blocking wavelengths below 540 nm were used to filter laser reflections. Slices of normalized radial velocity at fixed spanwise positions were presented to illustrate the

*normalized radial velocity at fixed spanwise locations for stereo reconstructed velocity field [61].*

*Experimental setup for PIV measurements performed at Purdue University (left) and sample results (right) of* 

*Sample PIV measurements in the diffuser passage of a high-speed centrifugal compressor at both the design* 

Conventionally, surface pressures are measured using hundreds of pressure taps or flush-mounted transducers to obtain a reasonable spatial distribution. This makes

development of the tip leakage flow across the rotor passage.

**3. Pressure-sensitive paints**

#### **Figure 9.** *Schematic of pressure-sensitive paint measurement system [36].*


#### **Table 4.**

*Representative studies that have used PSP technique in turbomachines.*

the measurements time-consuming and expensive. Recently, the introduction of pressure-sensitive paint (PSP) provides a new method for surface pressure measurement. Comparing to the conventional approaches by means of pressure taps or transducers which can only provide data at discrete points and are limited by installation locations, the PSP technique is very attractive; hence, it provides high-spatial-resolution pressure measurements without taps or transducers. The PSP technique is based on covering a surface with luminescent coatings. The luminescence of the coating is dependent on surface static pressure. With proper illumination, the surface pressure

**Figure 10.**

*Comparison of surface pressure distribution on a rotor suction side in a transonic axial compressor obtained using PSP (left) to CFD predictions (right) [68].*

distribution is obtained from images of illuminated surface. **Figure 9** shows all the essential optical and electrical components of a PSP system. It consists of various illumination devices, a local image and data-acquisition system, and an external calibration chamber.

The first aerodynamic study using PSP is performed by Pervushin et al. in 1985 to measure the pressure of air on the surface of wind tunnel models [62], and since then, numerous studies using PSP in external aerodynamics research have been conducted. However, unlike the well-established applications in external aerodynamic research, the application of PSP in turbomachines is quite limited. **Table 4** lists the studies in the open literature that have used pressuresensitive paint in turbomachines. A sample result of the PSP measurements together with the comparison to CFD results from the study conducted by Navarra et al. [68] is shown in **Figure 10**. The PSP measurement was conducted on the suction surface of the first-stage rotor of a state-of-the-art, full-scale transonic compressor.

## **4. Conclusions**

This chapter attempts to provide a comprehensive but brief summary of several advanced measurement techniques that have been used in turbomachines. For each measurement technique, the fundamental working principle was provided first and followed by discussion of its application in turbomachines. A list of representative research from the open literature was also provided for reference.

**37**

**Author details**

Fangyuan Lou

provided the original work is properly cited.

Purdue University, West Lafayette, USA

\*Address all correspondence to: louf@purdue.edu

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium,

## **Conflict of interest**

The author claims there is no conflict of interest.

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

*Rotating Machinery*

**Figure 10.**

calibration chamber.

transonic compressor.

**Conflict of interest**

**4. Conclusions**

*using PSP (left) to CFD predictions (right) [68].*

*Comparison of surface pressure distribution on a rotor suction side in a transonic axial compressor obtained* 

distribution is obtained from images of illuminated surface. **Figure 9** shows all the essential optical and electrical components of a PSP system. It consists of various illumination devices, a local image and data-acquisition system, and an external

The first aerodynamic study using PSP is performed by Pervushin et al. in 1985 to measure the pressure of air on the surface of wind tunnel models [62], and since then, numerous studies using PSP in external aerodynamics research have been conducted. However, unlike the well-established applications in external aerodynamic research, the application of PSP in turbomachines is quite limited. **Table 4** lists the studies in the open literature that have used pressuresensitive paint in turbomachines. A sample result of the PSP measurements together with the comparison to CFD results from the study conducted by Navarra et al. [68] is shown in **Figure 10**. The PSP measurement was conducted on the suction surface of the first-stage rotor of a state-of-the-art, full-scale

This chapter attempts to provide a comprehensive but brief summary of several advanced measurement techniques that have been used in turbomachines. For each measurement technique, the fundamental working principle was provided first and followed by discussion of its application in turbomachines. A list of representative

research from the open literature was also provided for reference.

The author claims there is no conflict of interest.

**36**

## **Author details**

Fangyuan Lou Purdue University, West Lafayette, USA

\*Address all correspondence to: louf@purdue.edu

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

## **References**

[1] Lakshminarayana B, Poncet A. A method of measuring threedimensional rotating wales behind turbomachinery rotors. Journal of Fluids Engineering. 1974;**96**(2):87-91

[2] Gorton CA, Lakshminarayana B. A method of measuring the three-dimensional mean flow and turbulence quantities inside a rotating turbo-machinery passage. Journal of Engineering for Power. 1976;**98**(2):137-144

[3] Hah C, Lakshminarayana B. Freestream turbulence effects on the development of a rotor wake. AIAA Journal. 1981;**19**(6):724-730

[4] Hodson HP, Huntsman I, Steele AB. An investigation of boundary layer development in a multistage LP turbine. Journal of Turbomachinery. 1994;**116**(3):375-383

[5] Camp TR, Shin HW. Turbulence intensity and length scale measurements in multistage compressors. Journal of Turbomachinery. 1995;**117**(1):38-46

[6] Witkowski AS, Chmielniak TJ, Strozik MD. Experimental study of a 3D wake decay and secondary flows behind a rotor blade row of a low speed compressor stage. In: ASME 1996 International Gas Turbine and Aeroengine Congress and Exhibition; 10 June 1996; American Society of Mechanical Engineers. 1996. p. V001T01A107

[7] Halstead DE, Wisler DC, Okiishi TH, Walker GJ, Hodson HP, Shin HW. Boundary layer development in axial compressors and turbines: Part 1 of 4—Composite picture. In: ASME 1995 International Gas Turbine and Aeroengine Congress and Exposition; 5 June 1995; American Society of Mechanical Engineers. 1995. p. V001T01A109

[8] Hsu ST, Wo AM. Near-wake measurement in a rotor/stator axial compressor using slanted hot-wire technique. Experiments in Fluids. 1997;**23**(5):441-444

[9] Ristic D, Lakshminarayana B. Threedimensional blade boundary layer and endwall flow development in the nozzle passage of a single stage turbine. Journal of Fluids Engineering. 1998;**120**(3):570-578

[10] Furukawa M, Saiki K, Nagayoshi K, Kuroumaru M, Inoue M. Effects of stream surface inclination on tip leakage flow fields in compressor rotors. In: ASME 1997 International Gas Turbine and Aeroengine Congress and Exhibition; 2 June 1997; American Society of Mechanical Engineers. 1997. p. V001T03A006

[11] Sentker A, Riess W. Experimental investigation of turbulent wake-blade interaction in axial compressors. International Journal of Heat and Fluid Flow. 2000;**21**(3):285-290

[12] Velarde-Suárez S, Ballesteros-Tajadura R, Santolaria-Morros C, González-Pérez J. Unsteady flow pattern characteristics downstream of a forwardcurved blades centrifugal fan. Journal of Fluids Engineering. 2001;**123**(2):265-270

[13] Pinarbasi A. Experimental hotwire measurements in a centrifugal compressor with vaned diffuser. International Journal of Heat and Fluid Flow. 2008;**29**(5):1512-1526

[14] Goodhand MN, Miller RJ. Compressor leading edge spikes: A new performance criterion. Journal of Turbomachinery. 2011;**133**(2):021006

[15] Weichert S, Day I. Detailed measurements of spike formation in an axial compressor. Journal of Turbomachinery. 2014;**136**(5):051006

**39**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

> [24] Abramian M, Howard JH. A rotating laser-Doppler anemometry system for unsteady relative flow measurements in model centrifugal impellers. Journal of Turbomachinery. 1994;**116**(2):260-268

[25] Zaccaria MA, Lakshminarayana B. Unsteady flow field due to nozzle wake interaction with the rotor in an axial flow turbine: Part I—Rotor passage flow field. In: ASME 1995 International Gas Turbine and Aeroengine Congress and Exposition; 5 June 1995; American Society of Mechanical Engineers. 1995.

[26] Adler D, Benyamin R. Experimental

propagation inside the flow passages of an axial gas turbine rotor. International Journal of Turbo and Jet-Engines.

[27] Ristic D, Lakshminarayana B, Chu S.

downstream of an axial-flow turbine rotor. Journal of Propulsion and Power.

[28] Faure TM, Michon GJ, Miton H, Vassilieff N. Laser Doppler anemometry measurements in an axial compressor stage. Journal of Propulsion and Power.

[29] Van Zante DE, To WM, Chen JP. Blade row interaction effects on the performance of a moderately loaded NASA transonic compressor stage. In: ASME Turbo Expo 2002: Power for Land, Sea, and Air; 1 January 2002; American Society of Mechanical Engineers. 2002. p. 969, 980

[30] Ibaraki S, Matsuo T, Kuma H, Sumida K, Suita T. Aerodynamics of a transonic centrifugal compressor impeller. Journal of Turbomachinery.

[31] Higashimori H, Hasagawa K, Sumida K, Suita T. Detailed flow study of Mach number 1.6 high transonic

2003;**125**(2):346-351

investigation of the stator wake

Three-dimensional flow field

p. V001T01A081

1999;**16**(4):193-206

1999;**15**(2):334-344

2001;**17**(3):481-491

[16] Wisler DC, Mossey PW. Gas velocity measurements within a compressor rotor passage using the laser Doppler velocimeter. Journal of Engineering for

[17] Pierzga MJ, Wood JR. Investigation of the three-dimensional flow field within a transonic fan rotor: Experiment and analysis. Journal of Engineering

[18] Strazisar AJ. Investigation of flow phenomena in a transonic fan rotor using laser anemometry. Journal of Engineering for Gas Turbines and Power. 1985;**107**(2):427-435

[19] Murthy KN, Lakshminarayana B. Laser Doppler velocimeter measurement in the tip region of a compressor rotor. AIAA Journal. 1986;**24**(5):807-814

[20] Beaudoin RJ, Miner SM, Flack RD. Laser velocimeter measurements in a centrifugal pump with a synchronously orbiting impeller. In: ASME 1990 International Gas Turbine and Aeroengine Congress and Exposition; 11 June 1990; American Society of Mechanical Engineers. 1990.

[21] Hathaway MD, Chriss RM, Wood JR,

[22] Farrell KJ, Billet ML. A correlation

of leakage vortex cavitation in axial-flow pumps. Journal of Fluids Engineering. 1994;**116**(3):551-557

[23] Fagan JR, Fleeter S. Comparison of optical measurement techniques for turbomachinery flow fields. Journal of Propulsion and Power. 1994;**10**(2):176-182

Strazisar AJ. Experimental and computational investigation of the NASA low-speed centrifugal compressor flow field. In: ASME 1992 International Gas Turbine and Aeroengine Congress and Exposition; 1 June 1992; American Society of Mechanical Engineers. 1992.

Power. 1973;**95**(2):91-96

for Gas Turbines and Power.

1985;**107**(2):436-448

p. V001T01A087

p. V001T01A074

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

[16] Wisler DC, Mossey PW. Gas velocity measurements within a compressor rotor passage using the laser Doppler velocimeter. Journal of Engineering for Power. 1973;**95**(2):91-96

[17] Pierzga MJ, Wood JR. Investigation of the three-dimensional flow field within a transonic fan rotor: Experiment and analysis. Journal of Engineering for Gas Turbines and Power. 1985;**107**(2):436-448

[18] Strazisar AJ. Investigation of flow phenomena in a transonic fan rotor using laser anemometry. Journal of Engineering for Gas Turbines and Power. 1985;**107**(2):427-435

[19] Murthy KN, Lakshminarayana B. Laser Doppler velocimeter measurement in the tip region of a compressor rotor. AIAA Journal. 1986;**24**(5):807-814

[20] Beaudoin RJ, Miner SM, Flack RD. Laser velocimeter measurements in a centrifugal pump with a synchronously orbiting impeller. In: ASME 1990 International Gas Turbine and Aeroengine Congress and Exposition; 11 June 1990; American Society of Mechanical Engineers. 1990. p. V001T01A087

[21] Hathaway MD, Chriss RM, Wood JR, Strazisar AJ. Experimental and computational investigation of the NASA low-speed centrifugal compressor flow field. In: ASME 1992 International Gas Turbine and Aeroengine Congress and Exposition; 1 June 1992; American Society of Mechanical Engineers. 1992. p. V001T01A074

[22] Farrell KJ, Billet ML. A correlation of leakage vortex cavitation in axial-flow pumps. Journal of Fluids Engineering. 1994;**116**(3):551-557

[23] Fagan JR, Fleeter S. Comparison of optical measurement techniques for turbomachinery flow fields. Journal of Propulsion and Power. 1994;**10**(2):176-182 [24] Abramian M, Howard JH. A rotating laser-Doppler anemometry system for unsteady relative flow measurements in model centrifugal impellers. Journal of Turbomachinery. 1994;**116**(2):260-268

[25] Zaccaria MA, Lakshminarayana B. Unsteady flow field due to nozzle wake interaction with the rotor in an axial flow turbine: Part I—Rotor passage flow field. In: ASME 1995 International Gas Turbine and Aeroengine Congress and Exposition; 5 June 1995; American Society of Mechanical Engineers. 1995. p. V001T01A081

[26] Adler D, Benyamin R. Experimental investigation of the stator wake propagation inside the flow passages of an axial gas turbine rotor. International Journal of Turbo and Jet-Engines. 1999;**16**(4):193-206

[27] Ristic D, Lakshminarayana B, Chu S. Three-dimensional flow field downstream of an axial-flow turbine rotor. Journal of Propulsion and Power. 1999;**15**(2):334-344

[28] Faure TM, Michon GJ, Miton H, Vassilieff N. Laser Doppler anemometry measurements in an axial compressor stage. Journal of Propulsion and Power. 2001;**17**(3):481-491

[29] Van Zante DE, To WM, Chen JP. Blade row interaction effects on the performance of a moderately loaded NASA transonic compressor stage. In: ASME Turbo Expo 2002: Power for Land, Sea, and Air; 1 January 2002; American Society of Mechanical Engineers. 2002. p. 969, 980

[30] Ibaraki S, Matsuo T, Kuma H, Sumida K, Suita T. Aerodynamics of a transonic centrifugal compressor impeller. Journal of Turbomachinery. 2003;**125**(2):346-351

[31] Higashimori H, Hasagawa K, Sumida K, Suita T. Detailed flow study of Mach number 1.6 high transonic

**38**

*Rotating Machinery*

[1] Lakshminarayana B, Poncet A. A method of measuring threedimensional rotating wales behind turbomachinery rotors. Journal of Fluids Engineering. 1974;**96**(2):87-91 [8] Hsu ST, Wo AM. Near-wake measurement in a rotor/stator axial compressor using slanted hot-wire technique. Experiments in Fluids.

[9] Ristic D, Lakshminarayana B. Threedimensional blade boundary layer and endwall flow development in the nozzle passage of a single stage turbine. Journal of Fluids Engineering.

[10] Furukawa M, Saiki K, Nagayoshi K, Kuroumaru M, Inoue M. Effects of stream surface inclination on tip leakage flow fields in compressor rotors. In: ASME 1997 International Gas Turbine and Aeroengine Congress and Exhibition; 2 June 1997; American Society of Mechanical Engineers. 1997.

[11] Sentker A, Riess W. Experimental investigation of turbulent wake-blade interaction in axial compressors. International Journal of Heat and Fluid

1997;**23**(5):441-444

1998;**120**(3):570-578

p. V001T03A006

Flow. 2000;**21**(3):285-290

[12] Velarde-Suárez S, Ballesteros-Tajadura R, Santolaria-Morros C, González-Pérez J. Unsteady flow pattern characteristics downstream of a forwardcurved blades centrifugal fan. Journal of Fluids Engineering. 2001;**123**(2):265-270

[13] Pinarbasi A. Experimental hotwire measurements in a centrifugal compressor with vaned diffuser. International Journal of Heat and Fluid

Flow. 2008;**29**(5):1512-1526

[14] Goodhand MN, Miller RJ. Compressor leading edge spikes: A new performance criterion. Journal of Turbomachinery. 2011;**133**(2):021006

[15] Weichert S, Day I. Detailed measurements of spike formation in an axial compressor. Journal of Turbomachinery. 2014;**136**(5):051006

[2] Gorton CA, Lakshminarayana B.

A method of measuring the three-dimensional mean flow and turbulence quantities inside a rotating turbo-machinery passage. Journal of Engineering for Power.

[3] Hah C, Lakshminarayana B. Freestream turbulence effects on the development of a rotor wake. AIAA

[4] Hodson HP, Huntsman I, Steele AB. An investigation of boundary layer development in a multistage LP turbine. Journal of Turbomachinery.

[5] Camp TR, Shin HW. Turbulence intensity and length scale measurements in multistage compressors. Journal of Turbomachinery. 1995;**117**(1):38-46

[6] Witkowski AS, Chmielniak TJ, Strozik MD. Experimental study of a 3D wake decay and secondary flows behind a rotor blade row of a low speed compressor stage. In: ASME 1996 International Gas Turbine and Aeroengine Congress and Exhibition; 10 June 1996; American Society of Mechanical Engineers. 1996.

[7] Halstead DE, Wisler DC, Okiishi TH, Walker GJ, Hodson HP, Shin HW. Boundary layer development in axial compressors and turbines: Part 1 of 4—Composite picture. In: ASME 1995 International Gas Turbine and Aeroengine Congress and Exposition; 5 June 1995; American Society of Mechanical Engineers. 1995.

Journal. 1981;**19**(6):724-730

1976;**98**(2):137-144

1994;**116**(3):375-383

p. V001T01A107

p. V001T01A109

**References**

flow with a shock wave in a pressure ratio 11 centrifugal compressor impeller. Journal of Turbomachinery. 2004;**126**(4):473-481

[32] Faure TM, Miton H, Vassilieff N. A laser Doppler anemometry technique for Reynolds stresses measurement. Experiments in Fluids. 2004;**37**(3):465-467

[33] Schleer M, Hong SS, Zangeneh M, Roduner C, Ribi B, Pløger F, et al. Investigation of an inversely designed centrifugal compressor stage—Part II: Experimental investigations. Journal of Turbomachinery. 2004;**126**(1):82-90

[34] Ibaraki S, Sumida K, Suita T. Design and off-design flow fields of a transonic centrifugal compressor impeller. In: ASME Turbo Expo 2009: Power for Land, Sea, and Air; 1 January 2009; American Society of Mechanical Engineers. 2009. pp. 1375-1384

[35] Gooding JW, Fabian JC, Key NL. LDV characterization of unsteady vaned diffuser flow in a centrifugal compressor. In: ASME Turbo Expo 2019: Turbomachinery Technical Conference & Exposition; 17 June 2019; American Society of Mechanical Engineers. 2019. p. GT2019-90476

[36] Tropea C, Yarin AL. Springer Handbook of Experimental Fluid Mechanics. Berlin, Heidelberg: Springer Science & Business Media; 2007

[37] Chow YC, Uzol O, Katz J. Flow nonuniformities and turbulent "hot spots" due to wake-blade and wake-wake interactions in a multistage turbomachine. Journal of Turbomachinery. 2002;**124**(4):553-563

[38] Paone N, Riethmuller ML, Van den Braembussche RA. Experimental investigation of the flow in the vaneless diffuser of a centrifugal pump by particle image displacement velocimetry. Experiments in Fluids. 1989;**7**(6):371-378

[39] Chu S, Dong R, Katz J. Relationship between unsteady flow, pressure fluctuations, and noise in a centrifugal pump—Part A: Use of PDV data to compute the pressure field. Journal of Fluids Engineering. 1995;**117**:24-29

[40] Chu S, Dong R, Katz J. Relationship between unsteady flow, pressure fluctuations, and noise in a centrifugal pump—Part B: Effects of bladetongue interactions. Journal of Fluids Engineering. 1995;**117**:30-35

[41] Day K, Lawless P, Fleeter S. Particle image velocimetry measurements in a low speed two stage research turbine. In: 32nd Joint Propulsion Conference and Exhibit. 1996. p. 2569

[42] Dong R, Chu S, Katz J. Effect of modification to tongue and impeller geometry on unsteady flow, pressure fluctuations, and noise in a centrifugal pump. Journal of Turbomachinery. 1997;**119**(3):506-515

[43] Wernet MP. Development of digital particle imaging velocimetry for use in turbomachinery. Experiments in Fluids. 2000;**28**(2):97-115

[44] Sinha M, Katz J. Quantitative visualization of the flow in a centrifugal pump with diffuser vanes—I: On flow structures and turbulence. Journal of Fluids Engineering. 2000;**122**(1):97-107

[45] Uzol O, Camci C. Aerodynamic loss characteristics of a turbine blade with trailing edge coolant ejection: Part 2— External aerodynamics, total pressure losses, and predictions. Journal of Turbomachinery. 2001;**123**(2):249-257

[46] Wernet MP, Bright MM, Skoch GJ. An investigation of surge in a highspeed centrifugal compressor using digital PIV. Journal of Turbomachinery. 2001;**123**(2):418-428

**41**

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

> [55] Ibaraki S, Matsuo T, Yokoyama T. Investigation of unsteady flow field in a vaned diffuser of a transonic centrifugal compressor. Journal of Turbomachinery.

> Copenhaver WW. PIV study of wake-rotor interactions in a transonic compressor at various operating conditions. Journal of Propulsion and Power. 2007;**23**(1):235-242

[57] Voges M, Beversdorff M, Willert C, Krain H. Application of particle image velocimetry to a transonic centrifugal compressor. Experiments in Fluids.

[58] Voges M, Willert CE, Mönig R, Müller MW, Schiffer HP. The challenge of stereo PIV measurements in the tip gap of a transonic compressor rotor with casing treatment. Experiments in

[59] Guillou E, Gancedo M, Gutmark E,

[60] Gancedo M, Gutmark E, Guillou E. Piv measurements of the flow at the inlet of a turbocharger centrifugal compressor with recirculation casing treatment near the inducer. Experiments

[61] Bhattacharya S, Berdanier RA, Vlachos PP, Key NL. A new particle image velocimetry technique for

[62] Ardasheva MM, Nevskii LB,

turbomachinery applications. Journal of Turbomachinery. 2016;**138**(12):124501

Pervushin GE. Measurement of pressure distribution by means of indicator coatings. Journal of Applied Mechanics and Technical Physics. 1985;**26**(4):469-474

[63] Sabroske K, Rabe D, Williams C. Pressure-sensitive paint investigation

2007;**129**(4):686-693

2007;**43**(2-3):371-384

Fluids. 2012;**52**(3):581-590

2012;**53**(3):619-635

in Fluids. 2016;**57**(2):16

Mohamed A. PIV investigation of the flow induced by a passive surge control method in a radial compressor. Experiments in Fluids.

[56] Estevadeordal J, Gorrell SE,

Meneveau C. Experimental investigation

[47] Uzol O, Chow YC, Katz J,

of unsteady flow field within a two-stage axial turbomachine using particle image velocimetry. Journal of Turbomachinery. 2002;**124**(4):542-552

[48] Sanders AJ, Papalia J, Fleeter S. Multi-blade row interactions in a transonic axial compressor: Part I—Stator particle image velocimetry (PIV) investigation. Journal of Turbomachinery. 2002;**124**(1):10-18

[49] Estevadeordal J, Gogineni S, Goss L, Copenhaver W, Gorrell S. Study of wake-blade interactions in a transonic compressor using flow visualization and DPIV. Journal of Fluids Engineering.

[50] Woisetschläger J, Mayrhofer N, Hampel B, Lang H, Sanz W. Laseroptical investigation of turbine wake flow. Experiments in Fluids.

[51] Uzol O, Chow YC, Katz J, Meneveau C. Average passage flow field and deterministic stresses in the tip and hub regions of a multistage turbomachine.

[52] Lee SJ, Paik BG, Yoon JH, Lee CM. Three-component velocity field measurements of propeller wake using a stereoscopic PIV technique. Experiments in Fluids.

2002;**124**(1):166-175

2003;**34**(3):371-378

2003;**125**(4):714-725

2004;**36**(4):575-585

2005;**39**(4):743-753

[53] Wernet MP, Van Zante D, Strazisar TJ, John WT, Prahst PS. Characterization of the tip clearance flow in an axial compressor using 3-D digital PIV. Experiments in Fluids.

[54] Yu XJ, Liu BJ. Stereoscopic PIV measurement of unsteady flows in an axial compressor stage. Experimental

Thermal and Fluid Science. 2007;**31**(8):1049-1060

Journal of Turbomachinery.

*Advance Measurement Techniques in Turbomachines DOI: http://dx.doi.org/10.5772/intechopen.85910*

*Rotating Machinery*

2004;**126**(4):473-481

2004;**37**(3):465-467

2004;**126**(1):82-90

p. GT2019-90476

flow with a shock wave in a pressure ratio 11 centrifugal compressor impeller. Journal of Turbomachinery. velocimetry. Experiments in Fluids.

[39] Chu S, Dong R, Katz J. Relationship between unsteady flow, pressure fluctuations, and noise in a centrifugal pump—Part A: Use of PDV data to compute the pressure field. Journal of Fluids Engineering. 1995;**117**:24-29

[40] Chu S, Dong R, Katz J. Relationship between unsteady flow, pressure fluctuations, and noise in a centrifugal pump—Part B: Effects of bladetongue interactions. Journal of Fluids

[41] Day K, Lawless P, Fleeter S. Particle image velocimetry measurements in a low speed two stage research turbine. In: 32nd Joint Propulsion Conference and

[42] Dong R, Chu S, Katz J. Effect of modification to tongue and impeller geometry on unsteady flow, pressure fluctuations, and noise in a centrifugal pump. Journal of Turbomachinery.

[43] Wernet MP. Development of digital particle imaging velocimetry for use in turbomachinery. Experiments in Fluids.

[45] Uzol O, Camci C. Aerodynamic loss characteristics of a turbine blade with trailing edge coolant ejection: Part 2— External aerodynamics, total pressure losses, and predictions. Journal of Turbomachinery. 2001;**123**(2):249-257

[46] Wernet MP, Bright MM, Skoch GJ. An investigation of surge in a highspeed centrifugal compressor using digital PIV. Journal of Turbomachinery.

2001;**123**(2):418-428

[44] Sinha M, Katz J. Quantitative visualization of the flow in a centrifugal pump with diffuser vanes—I: On flow structures and turbulence. Journal of Fluids Engineering. 2000;**122**(1):97-107

Engineering. 1995;**117**:30-35

Exhibit. 1996. p. 2569

1997;**119**(3):506-515

2000;**28**(2):97-115

1989;**7**(6):371-378

[32] Faure TM, Miton H, Vassilieff N.

[33] Schleer M, Hong SS, Zangeneh M, Roduner C, Ribi B, Pløger F, et al. Investigation of an inversely designed centrifugal compressor stage—Part II: Experimental investigations. Journal of Turbomachinery.

[34] Ibaraki S, Sumida K, Suita T. Design and off-design flow fields of a transonic centrifugal compressor impeller. In: ASME Turbo Expo 2009: Power for Land, Sea, and Air; 1 January 2009; American Society of Mechanical Engineers. 2009. pp. 1375-1384

[35] Gooding JW, Fabian JC, Key NL. LDV characterization of unsteady vaned diffuser flow in a centrifugal compressor. In: ASME Turbo Expo 2019: Turbomachinery Technical Conference & Exposition; 17 June 2019; American Society of Mechanical Engineers. 2019.

[36] Tropea C, Yarin AL. Springer Handbook of Experimental Fluid Mechanics. Berlin, Heidelberg: Springer

Science & Business Media; 2007

[37] Chow YC, Uzol O, Katz J. Flow nonuniformities and turbulent "hot spots" due to wake-blade and wake-wake interactions in a multistage turbomachine. Journal of Turbomachinery. 2002;**124**(4):553-563

[38] Paone N, Riethmuller ML, Van den Braembussche RA. Experimental investigation of the flow in the vaneless diffuser of a centrifugal pump by particle image displacement

A laser Doppler anemometry technique for Reynolds stresses measurement. Experiments in Fluids.

**40**

[47] Uzol O, Chow YC, Katz J, Meneveau C. Experimental investigation of unsteady flow field within a two-stage axial turbomachine using particle image velocimetry. Journal of Turbomachinery. 2002;**124**(4):542-552

[48] Sanders AJ, Papalia J, Fleeter S. Multi-blade row interactions in a transonic axial compressor: Part I—Stator particle image velocimetry (PIV) investigation. Journal of Turbomachinery. 2002;**124**(1):10-18

[49] Estevadeordal J, Gogineni S, Goss L, Copenhaver W, Gorrell S. Study of wake-blade interactions in a transonic compressor using flow visualization and DPIV. Journal of Fluids Engineering. 2002;**124**(1):166-175

[50] Woisetschläger J, Mayrhofer N, Hampel B, Lang H, Sanz W. Laseroptical investigation of turbine wake flow. Experiments in Fluids. 2003;**34**(3):371-378

[51] Uzol O, Chow YC, Katz J, Meneveau C. Average passage flow field and deterministic stresses in the tip and hub regions of a multistage turbomachine. Journal of Turbomachinery. 2003;**125**(4):714-725

[52] Lee SJ, Paik BG, Yoon JH, Lee CM. Three-component velocity field measurements of propeller wake using a stereoscopic PIV technique. Experiments in Fluids. 2004;**36**(4):575-585

[53] Wernet MP, Van Zante D, Strazisar TJ, John WT, Prahst PS. Characterization of the tip clearance flow in an axial compressor using 3-D digital PIV. Experiments in Fluids. 2005;**39**(4):743-753

[54] Yu XJ, Liu BJ. Stereoscopic PIV measurement of unsteady flows in an axial compressor stage. Experimental Thermal and Fluid Science. 2007;**31**(8):1049-1060

[55] Ibaraki S, Matsuo T, Yokoyama T. Investigation of unsteady flow field in a vaned diffuser of a transonic centrifugal compressor. Journal of Turbomachinery. 2007;**129**(4):686-693

[56] Estevadeordal J, Gorrell SE, Copenhaver WW. PIV study of wake-rotor interactions in a transonic compressor at various operating conditions. Journal of Propulsion and Power. 2007;**23**(1):235-242

[57] Voges M, Beversdorff M, Willert C, Krain H. Application of particle image velocimetry to a transonic centrifugal compressor. Experiments in Fluids. 2007;**43**(2-3):371-384

[58] Voges M, Willert CE, Mönig R, Müller MW, Schiffer HP. The challenge of stereo PIV measurements in the tip gap of a transonic compressor rotor with casing treatment. Experiments in Fluids. 2012;**52**(3):581-590

[59] Guillou E, Gancedo M, Gutmark E, Mohamed A. PIV investigation of the flow induced by a passive surge control method in a radial compressor. Experiments in Fluids. 2012;**53**(3):619-635

[60] Gancedo M, Gutmark E, Guillou E. Piv measurements of the flow at the inlet of a turbocharger centrifugal compressor with recirculation casing treatment near the inducer. Experiments in Fluids. 2016;**57**(2):16

[61] Bhattacharya S, Berdanier RA, Vlachos PP, Key NL. A new particle image velocimetry technique for turbomachinery applications. Journal of Turbomachinery. 2016;**138**(12):124501

[62] Ardasheva MM, Nevskii LB, Pervushin GE. Measurement of pressure distribution by means of indicator coatings. Journal of Applied Mechanics and Technical Physics. 1985;**26**(4):469-474

[63] Sabroske K, Rabe D, Williams C. Pressure-sensitive paint investigation for turbomachinery application. In: ASME 1995 International Gas Turbine and Aeroengine Congress and Exposition; 5 June 1995; American Society of Mechanical Engineers. 1995. p. V001T01A021

[64] Liu T, Torgerson S, Sullivan J, Johnston R, Fleeter S, Liu T, et al. Rotor blade pressure measurement in a high speed axial compressor using pressure and temperature sensitive paints. In: 35th Aerospace Sciences Meeting and Exhibit. 1997. p. 162

[65] Navarra KR. Development of the pressure-sensitive-paint technique for advanced turbomachinery applications [master thesis]. Virginia Tech;

[66] Bencic T. Rotating pressure and temperature measurements on scalemodel fans using luminescent paints. In: 34th AIAA/ASME/SAE/ASEE Joint Propulsion Conference and Exhibit; 1 January 1998. p. 3452

[67] Engler RH, Klein C, Trinks O. Pressure sensitive paint systems for pressure distribution measurements in wind tunnels and turbomachines. Measurement Science and Technology. 2000;**11**(7):1077

[68] Navarra KR, Rabe DC, Fonov SD, Goss LP, Hah C. The application of pressure-and temperature-sensitive paints to an advanced compressor. Journal of Turbomachinery. 2001;**123**(4):823-829

[69] Gregory J, Sakaue H, Sullivan J. Unsteady pressure measurements in turbomachinery using porous pressure sensitive paint. In: 40th AIAA Aerospace Sciences Meeting & Exhibit. 2002. p. 84

[70] Lepicovsky J, Bencic T. Use of pressure-sensitive paint for diagnostics in turbomachinery flows with shocks. Experiments in Fluids. 2002;**33**(4):531-538

[71] Gregory J. Porous pressure-sensitive paint for measurement of unsteady pressures in turbomachinery. In: 42nd AIAA Aerospace Sciences Meeting and Exhibit; 5 January 2004. 2004. p. 294

[72] Suryanarayanan A, Ozturk B, Schobeiri MT, Han JC. Film-cooling effectiveness on a rotating turbine platform using pressure sensitive paint technique. Journal of Turbomachinery. 2010;**132**(4):041001

[73] Narzary DP, Liu KC, Rallabandi AP, Han JC. Influence of coolant density on turbine blade film-cooling using pressure sensitive paint technique. Journal of Turbomachinery. 2012;**134**(3):031006

**43**

**1. Introduction**

**Chapter 4**

*Akanksha Singh*

**Abstract**

Development and Control of

Direct-Drive Wind Turbines

turbine topology is also presented in this chapter.

**Keywords:** boost CSI, direct-drive wind turbine, finite element analysis, permanent magnet synchronous generator, wind turbine topologies

Wind power is one of the fastest growing energies and the global capacity has increased to 433 GW by the end of 2015 [1, 2]. The two most commonly used topologies for wind turbines are based on the Doubly Fed Induction Generator (DFIG) or Permanent Magnet Synchronous Generator (PMSG) [3]. In the DFIG configuration, the stator of the generator is directly connected to the grid while the rotor does not require to rotate at the fixed synchronous speed. The rotor is connected to the turbine shaft through a gearbox. In the PMSG-based wind turbines, the speed of the rotating magnetic field and the rotor is the same and therefore, it is connected to the grid through power converters. The PMSG-based wind turbines require a power electronics interface to be connected to the grid which provides the flexibility of using these turbines with or without gearbox between the turbine shaft and the generator shaft [3]. It has been shown that the gearboxes cause more

Generator-Converter Topology for

In this chapter, a new topology for Direct-Drive Wind Turbines (DDWTs) with a low-voltage generator design is presented in order to eliminate the required dc-bus capacitors or dc-link inductors. In the presented topology, the grid-side converter is replaced by a boost Current Source Inverter (CSI) therefore removing the need for the dc-bus electrolytic capacitors which results in increasing the system lifetime. In the developed topology, the synchronous inductance of the generator is utilized. This facilitates the elimination of the intrinsically required dc-link inductor in the CSI which further contributes to a reduction in the overall system weight and size. The boost CSI is capable of converting a low dc voltage to a higher line-to-line voltage. This results in the implementation of a low-voltage generator for DDWTs. The feasibility of the presented low-voltage generator is investigated through Finite Element (FE) computations. In this chapter, a modified 1.5 MW low-voltage generator for the proposed topology is compared with an existing 1.5 MW Permanent Magnet (PM) synchronous generator for DDWTs. The feasibility of the presented topology of generator-converter for DDWTs is verified through simulations and laboratory tests. Furthermore, the controls developed for the developed wind

## **Chapter 4**

*Rotating Machinery*

p. V001T01A021

Exhibit. 1997. p. 162

January 1998. p. 3452

2000;**11**(7):1077

2001;**123**(4):823-829

2002. p. 84

2002;**33**(4):531-538

for turbomachinery application. In: ASME 1995 International Gas Turbine and Aeroengine Congress and Exposition; 5 June 1995; American Society of Mechanical Engineers. 1995. [71] Gregory J. Porous pressure-sensitive paint for measurement of unsteady pressures in turbomachinery. In: 42nd AIAA Aerospace Sciences Meeting and Exhibit; 5 January 2004. 2004. p. 294

[72] Suryanarayanan A, Ozturk B, Schobeiri MT, Han JC. Film-cooling effectiveness on a rotating turbine platform using pressure sensitive paint technique. Journal of Turbomachinery.

[73] Narzary DP, Liu KC, Rallabandi AP, Han JC. Influence of coolant density on turbine blade film-cooling using pressure sensitive paint technique. Journal of Turbomachinery.

2010;**132**(4):041001

2012;**134**(3):031006

[64] Liu T, Torgerson S, Sullivan J, Johnston R, Fleeter S, Liu T, et al. Rotor blade pressure measurement in a high speed axial compressor using pressure and temperature sensitive paints. In: 35th Aerospace Sciences Meeting and

[65] Navarra KR. Development of the pressure-sensitive-paint technique for advanced turbomachinery applications

[66] Bencic T. Rotating pressure and temperature measurements on scalemodel fans using luminescent paints. In: 34th AIAA/ASME/SAE/ASEE Joint Propulsion Conference and Exhibit; 1

[67] Engler RH, Klein C, Trinks O. Pressure sensitive paint systems for pressure distribution measurements in wind tunnels and turbomachines. Measurement Science and Technology.

[68] Navarra KR, Rabe DC, Fonov SD, Goss LP, Hah C. The application of pressure-and temperature-sensitive paints to an advanced compressor. Journal of Turbomachinery.

[69] Gregory J, Sakaue H, Sullivan J. Unsteady pressure measurements in turbomachinery using porous pressure sensitive paint. In: 40th AIAA Aerospace Sciences Meeting & Exhibit.

[70] Lepicovsky J, Bencic T. Use of pressure-sensitive paint for diagnostics in turbomachinery flows with shocks. Experiments in Fluids.

[master thesis]. Virginia Tech;

**42**

## Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines

*Akanksha Singh*

## **Abstract**

In this chapter, a new topology for Direct-Drive Wind Turbines (DDWTs) with a low-voltage generator design is presented in order to eliminate the required dc-bus capacitors or dc-link inductors. In the presented topology, the grid-side converter is replaced by a boost Current Source Inverter (CSI) therefore removing the need for the dc-bus electrolytic capacitors which results in increasing the system lifetime. In the developed topology, the synchronous inductance of the generator is utilized. This facilitates the elimination of the intrinsically required dc-link inductor in the CSI which further contributes to a reduction in the overall system weight and size. The boost CSI is capable of converting a low dc voltage to a higher line-to-line voltage. This results in the implementation of a low-voltage generator for DDWTs. The feasibility of the presented low-voltage generator is investigated through Finite Element (FE) computations. In this chapter, a modified 1.5 MW low-voltage generator for the proposed topology is compared with an existing 1.5 MW Permanent Magnet (PM) synchronous generator for DDWTs. The feasibility of the presented topology of generator-converter for DDWTs is verified through simulations and laboratory tests. Furthermore, the controls developed for the developed wind turbine topology is also presented in this chapter.

**Keywords:** boost CSI, direct-drive wind turbine, finite element analysis, permanent magnet synchronous generator, wind turbine topologies

## **1. Introduction**

Wind power is one of the fastest growing energies and the global capacity has increased to 433 GW by the end of 2015 [1, 2]. The two most commonly used topologies for wind turbines are based on the Doubly Fed Induction Generator (DFIG) or Permanent Magnet Synchronous Generator (PMSG) [3]. In the DFIG configuration, the stator of the generator is directly connected to the grid while the rotor does not require to rotate at the fixed synchronous speed. The rotor is connected to the turbine shaft through a gearbox. In the PMSG-based wind turbines, the speed of the rotating magnetic field and the rotor is the same and therefore, it is connected to the grid through power converters. The PMSG-based wind turbines require a power electronics interface to be connected to the grid which provides the flexibility of using these turbines with or without gearbox between the turbine shaft and the generator shaft [3]. It has been shown that the gearboxes cause more

downtime than any other component in a wind turbine [3–5]. It is worth noting that the gearboxes are responsible for 10% of the wind turbine failures which result in about 20% of the total wind turbine downtime [4–6]. Recent investigations reveal that gearboxes in wind turbines, which were supposed to last 20 years, might fail in 7–10 years [7, 8]. The Direct-Drive Wind Turbines (DDWTs) do not have a gearbox between the turbine rotor and the generator shaft. There is a definite trend toward DDWTs as is predicted in the research papers and trade articles but there are some major concerns that must be overcome in order to achieve higher market penetration [1, 9]. The power electronics interface in a DDWT is rated for full power transfer and it consists of an active rectifier and a grid-side converter, connected through a capacitor bank forming the dc-bus. The power electronics interface is one of the most vulnerable components in wind turbines [4, 5, 7]. A significant percentage of these failures have been attributed to the dc-bus electrolytic capacitors [10, 11]. Many methods which try to determine the remaining lifespan of electrolytic capacitors exist and are utilized for scheduled maintenance planning [5]. These methods do not serve to extend the life of the capacitors and contribute to an increase in the system downtime [7]. The other major concern which prevents the proliferation of DDWTs is the large size of PMSGs [12, 13]. The large size of the generator is a result of the generator shaft rotating at the same speed as the turbine rotor shaft. The low angular speed in PMSGs increases the required pole surface and the number of poles, which results in a high cost and volume of the generator.

In the most commonly used DDWT configuration the power electronics interface comprises of back-to-back Voltage Source Converters (VSCs) connected through electrolytic capacitors [14]. In this paper, the grid-side VSC is replaced by the boost Current Source Inverter (CSI). The proposed boost CSI topology of the converter results in the elimination of the dc-bus electrolytic capacitors, which are one of the frequently failing components in the existing topologies of direct-drive wind turbines without any supplementary component at the dc-bus [10, 11, 15]. The boost CSI converts the low dc voltage of the generator rectifier to the higher voltage level which facilitates the implementation of low voltage PMSG with lower weight and volume [16]. In addition to the introduction, this chapter has five more sections. In Section 2, the existing wind turbine topologies are reviewed, and the topology of the developed system is introduced. Section 3 presents the feasibility of a 1.5 MW low-voltage PMSG for the presented system. This section also presents a comparison between the generator for the developed system and an existing 1.5 MW PMSG for a DDWT using Finite Element (FE) computations. The controls developed for the new DDWT topology are presented in Section 4. The simulation results demonstrating the feasibility of the developed system are presented in Section 5. A laboratory scale setup is used to experimentally confirm the feasibility of the proposed topology in Section 6. Section 7 presents the conclusions of the chapter.

## **2. Wind turbine topologies**

In this section, several wind turbine topologies are reviewed prior to introducing the inductorless current source generator-converter topology.

#### **2.1 Indirect drive wind turbine topologies**

In the majority of wind turbine topologies, the windmill rotor shaft is connected to the generator through a gearbox [14, 17]. The indirect drive wind turbines can be assembled using various types of generators and converters. Most commonly

**45**

**Figure 1.**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

used indirect wind turbine utilizes the DFIG configuration of the drivetrain. In this configuration, the rotor is connected to the grid through a back-to-back converter while the stator windings are directly connected to the grid and. This back-to-back converter includes of two bidirectional converters connected at the dc-bus, which formed by a capacitor bank. In this configuration, the major part of the power injected into the grid is through the stator of the generator and the back-to-back converters transfer only a fraction of the power produced by the wind turbine into the grid. The PMSG with the same back-to-back converter topology is another common indirect drive wind turbine topology. Unlike in the DFIG configuration, the

The DDWT is a relatively new topology and has the turbine rotor shaft directly connected to the generator shaft without any gearbox. **Figure 1** shows commonly used topologies for DDWTs using the PMSG. **Figure 1(a)** displays the power electronics interface which consists of a diode bridge rectifier, dc-dc boost converter, and grid-tied VSC [17, 18]. **Figure 1(b)** shows another DDWT topology with power electronics interface consisting of grid-tied back-to-back Current Source Converters (CSCs) connected through a dc-link inductor [19, 20]. **Figure 1(c)** demonstrates the most frequently used configuration of DDWTs, comprising of a PMSG connected

*Common DDWT topologies with (a) PMSG connected to diode-bridge, boost converter and VSI, (b) PMSG* 

*connected to back-to-back CSCs, and (c) PMSG connected to back-to-back VSCs.*

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

converter is rated for the full power.

**2.2 Direct drive wind turbine (DDWT) topologies**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

used indirect wind turbine utilizes the DFIG configuration of the drivetrain. In this configuration, the rotor is connected to the grid through a back-to-back converter while the stator windings are directly connected to the grid and. This back-to-back converter includes of two bidirectional converters connected at the dc-bus, which formed by a capacitor bank. In this configuration, the major part of the power injected into the grid is through the stator of the generator and the back-to-back converters transfer only a fraction of the power produced by the wind turbine into the grid. The PMSG with the same back-to-back converter topology is another common indirect drive wind turbine topology. Unlike in the DFIG configuration, the converter is rated for the full power.

## **2.2 Direct drive wind turbine (DDWT) topologies**

The DDWT is a relatively new topology and has the turbine rotor shaft directly connected to the generator shaft without any gearbox. **Figure 1** shows commonly used topologies for DDWTs using the PMSG. **Figure 1(a)** displays the power electronics interface which consists of a diode bridge rectifier, dc-dc boost converter, and grid-tied VSC [17, 18]. **Figure 1(b)** shows another DDWT topology with power electronics interface consisting of grid-tied back-to-back Current Source Converters (CSCs) connected through a dc-link inductor [19, 20]. **Figure 1(c)** demonstrates the most frequently used configuration of DDWTs, comprising of a PMSG connected

#### **Figure 1.**

*Common DDWT topologies with (a) PMSG connected to diode-bridge, boost converter and VSI, (b) PMSG connected to back-to-back CSCs, and (c) PMSG connected to back-to-back VSCs.*

*Rotating Machinery*

downtime than any other component in a wind turbine [3–5]. It is worth noting that the gearboxes are responsible for 10% of the wind turbine failures which result in about 20% of the total wind turbine downtime [4–6]. Recent investigations reveal that gearboxes in wind turbines, which were supposed to last 20 years, might fail in 7–10 years [7, 8]. The Direct-Drive Wind Turbines (DDWTs) do not have a gearbox between the turbine rotor and the generator shaft. There is a definite trend toward DDWTs as is predicted in the research papers and trade articles but there are some major concerns that must be overcome in order to achieve higher market penetration [1, 9]. The power electronics interface in a DDWT is rated for full power transfer and it consists of an active rectifier and a grid-side converter, connected through a capacitor bank forming the dc-bus. The power electronics interface is one of the most vulnerable components in wind turbines [4, 5, 7]. A significant percentage of these failures have been attributed to the dc-bus electrolytic capacitors [10, 11]. Many methods which try to determine the remaining lifespan of electrolytic capacitors exist and are utilized for scheduled maintenance planning [5]. These methods do not serve to extend the life of the capacitors and contribute to an increase in the system downtime [7]. The other major concern which prevents the proliferation of DDWTs is the large size of PMSGs [12, 13]. The large size of the generator is a result of the generator shaft rotating at the same speed as the turbine rotor shaft. The low angular speed in PMSGs increases the required pole surface and the number of

poles, which results in a high cost and volume of the generator.

presents the conclusions of the chapter.

**2.1 Indirect drive wind turbine topologies**

**2. Wind turbine topologies**

In the most commonly used DDWT configuration the power electronics interface comprises of back-to-back Voltage Source Converters (VSCs) connected through electrolytic capacitors [14]. In this paper, the grid-side VSC is replaced by the boost Current Source Inverter (CSI). The proposed boost CSI topology of the converter results in the elimination of the dc-bus electrolytic capacitors, which are one of the frequently failing components in the existing topologies of direct-drive wind turbines without any supplementary component at the dc-bus [10, 11, 15]. The boost CSI converts the low dc voltage of the generator rectifier to the higher voltage level which facilitates the implementation of low voltage PMSG with lower weight and volume [16]. In addition to the introduction, this chapter has five more sections. In Section 2, the existing wind turbine topologies are reviewed, and the topology of the developed system is introduced. Section 3 presents the feasibility of a 1.5 MW low-voltage PMSG for the presented system. This section also presents a comparison between the generator for the developed system and an existing 1.5 MW PMSG for a DDWT using Finite Element (FE) computations. The controls developed for the new DDWT topology are presented in Section 4. The simulation results demonstrating the feasibility of the developed system are presented in Section 5. A laboratory scale setup is used to experimentally confirm the feasibility of the proposed topology in Section 6. Section 7

In this section, several wind turbine topologies are reviewed prior to introducing

In the majority of wind turbine topologies, the windmill rotor shaft is connected to the generator through a gearbox [14, 17]. The indirect drive wind turbines can be assembled using various types of generators and converters. Most commonly

the inductorless current source generator-converter topology.

**44**

to the grid through two VSCs [17, 21, 22]. The dc-bus between the VSCs is formed using electrolytic capacitors in order to regulate and stabilize the dc-bus voltage. In some cases, multilevel converters are used to form the power electronics interface in the drivetrain. The topology of the multilevel converters can be an H-bridge backto-back converter or a neutral-point-clamped back-to-back converter [17, 22]. Most wind turbine topologies either have a dc-bus formed by electrolytic capacitors or a dc-link formed by inductors. The electrolytic capacitors are one of the most failureprone components that adversely impact the system reliability [5]. The failure of these capacitors has a significant impact on the maintenance cost especially in the case of offshore wind turbines [13, 23]. The dc-link inductor is bulky and adds to the system loss reducing the system's efficiency [19]. In some cases, the grid side converter is located away from the wind turbine and the transmission cable is used to realize the dc-link inductor [20]. Additionally, the lack of gearbox in the DDWT increases the size of the PMSG and the capital cost of the overall system [13, 23–25]. In the following, the proposed DDWT topology is described.

## **2.3 Proposed DDWT topology**

The proposed topology for the DDWT is presented in **Figure 2**. In this subsection, different parts of the proposed topology are explained. First, the power electronics interface topology is presented, and then, the PMSG design flexibility provided by implementing the boost CSI is described.

## *2.3.1 The power electronics interface*

The power electronics interface in the proposed topology is a back-to-back converter made up of a three-phase VSC and a boost CSI, as shown in **Figure 2**. The boost CSI is equipped by Reverse Blocking IGBTs (RB-IGBT), and no dc-bus capacitor or dc-link inductor exists in the proposed DDWT. As a result of avoiding the dc-bus capacitor, the system mean-time-between-failure can significantly be improved in comparison with the conventional VSC-based systems [26]. While any CSI requires an inductor at the dc-link, the proposed DDWT eliminates the dc-link inductor by utilizing the generator synchronous inductance, *Ls* [15]. In order to achieve a Total Harmonic Distortion (THD) of the current waveforms at the inverter output acceptable under the IEEE 1547–2018 interconnection standards, the operation of the boost CSI must always be in Continuous Conduction Mode (CCM). The minimum dc-link inductance, *Ldc* required to keep the inverter in CCM has been derived in [27]. However, the synchronous inductance, *Ls*, will be less than the *Ldc*, i.e. *Ls* = √ \_\_\_ 3/2 *Ldc*. This is because Thevenin's equivalent inductance of the generator-converter from the dc-bus is almost (3/2)*Ls* and *Irms* ≈ √ \_\_\_ 2/3 *Idc*. Also, a traditional CSI operated using space vector PWM switching provides a maximum

#### **Figure 2.**

*The proposed wind turbine system topology with PMSG connected to a VSC and the boost inverter equipped by RB-IGBTs.*

**47**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

boost ratio, *VLLrms*/*Vdc*, of about 1.2 [28]. However, the boost CSI is modulated using the Phasor Pulse Width Modulation (PPWM) switching technique, providing a boost ratio more than three [29–31]. As a result, a low voltage PMSG design will be possible for the proposed DDWT. The advantages gained by using a low voltage

The low dc-side voltage required by the boost CSI enables a low output ac-voltage from the PMSG. In a traditional DDWT topology, the voltage at the generator and the grid should have almost the same rms value [32, 33]. However, if the developed topology is utilized, then, for example, 120*VLLrms* from the PM generator is sufficient to generate an rms line-to-line voltage of 480 V at the converter output. As described in [15], consider the emf equation *E* ∝ *Nphp*ω*<sup>m</sup>* ϕ*<sup>p</sup>* , where, *E* is the peak induced voltage per phase, *Nph* is the number of winding turns per phase, ω*m* is the mechanical angular velocity, *p* is the total number of poles, and ϕ*p* is the maximum magnetic flux per pole, [33]. Here ω*m* is dependent on the wind speed, and the fundamental magnetic flux per pole is obtained from ϕ*<sup>p</sup>* ≅ (4/*π*)*Bm*(2*π*/*p*)(*D*/2)*l* = 4*DlBm*/*p*, where *D* is the mid-airgap diameter, *Bm* is the flux density, and *l* is the stack length. Accordingly, for a given angular velocity, permanent magnet material, and air-gap height, *lg* one can say *E* ∝ *NphDl*, which means that a generator with a lower output voltage requires a lower value of *NphDl*. On the other hand, the minimum synchronous inductance, required for the boost CSI to operate in CCM, restricts the maxi-

2

)

In this section, a 1.5 MW PMSG is designed for the proposed DDWT topology

The generator is designed using GenAC toolbox of MagneForce Finite Element (FE) software. The designed machine is a low-speed, concentrated overlapping tooth coil double layered winding, three-phase PMSG rated at 1.5 MW. The generator is designed to be radial flux, cylindrical rotor, rotor surface mounted pole machine so that it can be compared with the existing 1.5 MW PMSG [34]. For the sake of comparison, both the existing and new 1.5 MW systems have the rated speed of 19.65 rpm, and the same generator airgap of 0.6 mm. Also, the line-to-line voltage of the new generator is 450 V, while the existing generator line-to-line volt-

**Figure 3** shows the cross-sectional view of the new generator, including the stator slot and tooth dimensions in details. Moreover, the insulation thickness in each slot is 3.5 mm. The overall slot-fill is about 62%. Also, the rotor shaft bore diameter is 700 mm. The stator bore diameter is 3500 mm, the stator length is 600 mm, and the airgap length, *lg*, is 0.6 mm. The PM material is sintered Neodymium Iron Boron (NdFeB) with the maximum residual flux density of *Br* = 1.2 T, and maximum coercive force of 24 k Oersted (i.e. *Hc* =1910 kA/m). The permanent magnet pole

demonstrates the electromagnetic flux and the flux density distribution over the

with a pole arc to pole pitch ratio of 0.8. **Figure 4**

[33]. Hence, the desired values of

1/2 ≅ *E* and synchronous inductance, *Ls*,

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

generator are discussed in the next subsection.

mum number of poles, since *Ls* ∝ (*Dl*/*lg*)(*Nph*/*p*)

are herein used as the PMSG design inputs for a given rated power.

and then the design is compared with an existing 1.5 MW PMSG [34].

the generator output voltage, (*E*<sup>2</sup> − (ω*sLsI*)<sup>2</sup>

**3. Generator design and comparison**

**3.1 New generator design**

age is 770 V, see **Table 1**.

dimension is as 260 × 100mm<sup>2</sup>

*2.3.2 The permanent magnet generator*

#### *Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

boost ratio, *VLLrms*/*Vdc*, of about 1.2 [28]. However, the boost CSI is modulated using the Phasor Pulse Width Modulation (PPWM) switching technique, providing a boost ratio more than three [29–31]. As a result, a low voltage PMSG design will be possible for the proposed DDWT. The advantages gained by using a low voltage generator are discussed in the next subsection.

### *2.3.2 The permanent magnet generator*

*Rotating Machinery*

to the grid through two VSCs [17, 21, 22]. The dc-bus between the VSCs is formed using electrolytic capacitors in order to regulate and stabilize the dc-bus voltage. In some cases, multilevel converters are used to form the power electronics interface in the drivetrain. The topology of the multilevel converters can be an H-bridge backto-back converter or a neutral-point-clamped back-to-back converter [17, 22]. Most wind turbine topologies either have a dc-bus formed by electrolytic capacitors or a dc-link formed by inductors. The electrolytic capacitors are one of the most failureprone components that adversely impact the system reliability [5]. The failure of these capacitors has a significant impact on the maintenance cost especially in the case of offshore wind turbines [13, 23]. The dc-link inductor is bulky and adds to the system loss reducing the system's efficiency [19]. In some cases, the grid side converter is located away from the wind turbine and the transmission cable is used to realize the dc-link inductor [20]. Additionally, the lack of gearbox in the DDWT increases the size of the PMSG and the capital cost of the overall system [13, 23–25].

The proposed topology for the DDWT is presented in **Figure 2**. In this subsec-

The power electronics interface in the proposed topology is a back-to-back converter made up of a three-phase VSC and a boost CSI, as shown in **Figure 2**. The boost CSI is equipped by Reverse Blocking IGBTs (RB-IGBT), and no dc-bus capacitor or dc-link inductor exists in the proposed DDWT. As a result of avoiding the dc-bus capacitor, the system mean-time-between-failure can significantly be improved in comparison with the conventional VSC-based systems [26]. While any CSI requires an inductor at the dc-link, the proposed DDWT eliminates the dc-link inductor by utilizing the generator synchronous inductance, *Ls* [15]. In order to achieve a Total Harmonic Distortion (THD) of the current waveforms at the inverter output acceptable under the IEEE 1547–2018 interconnection standards, the operation of the boost CSI must always be in Continuous Conduction Mode (CCM). The minimum dc-link inductance, *Ldc* required to keep the inverter in CCM has been derived in [27]. However, the synchronous inductance, *Ls*, will be less than

traditional CSI operated using space vector PWM switching provides a maximum

*The proposed wind turbine system topology with PMSG connected to a VSC and the boost inverter equipped by* 

3/2 *Ldc*. This is because Thevenin's equivalent inductance of the

\_\_\_

2/3 *Idc*. Also, a

tion, different parts of the proposed topology are explained. First, the power electronics interface topology is presented, and then, the PMSG design flexibility

In the following, the proposed DDWT topology is described.

provided by implementing the boost CSI is described.

**2.3 Proposed DDWT topology**

*2.3.1 The power electronics interface*

\_\_\_

generator-converter from the dc-bus is almost (3/2)*Ls* and *Irms* ≈ √

**46**

**Figure 2.**

*RB-IGBTs.*

the *Ldc*, i.e. *Ls* = √

The low dc-side voltage required by the boost CSI enables a low output ac-voltage from the PMSG. In a traditional DDWT topology, the voltage at the generator and the grid should have almost the same rms value [32, 33]. However, if the developed topology is utilized, then, for example, 120*VLLrms* from the PM generator is sufficient to generate an rms line-to-line voltage of 480 V at the converter output. As described in [15], consider the emf equation *E* ∝ *Nphp*ω*<sup>m</sup>* ϕ*<sup>p</sup>* , where, *E* is the peak induced voltage per phase, *Nph* is the number of winding turns per phase, ω*m* is the mechanical angular velocity, *p* is the total number of poles, and ϕ*p* is the maximum magnetic flux per pole, [33]. Here ω*m* is dependent on the wind speed, and the fundamental magnetic flux per pole is obtained from ϕ*<sup>p</sup>* ≅ (4/*π*)*Bm*(2*π*/*p*)(*D*/2)*l* = 4*DlBm*/*p*, where *D* is the mid-airgap diameter, *Bm* is the flux density, and *l* is the stack length. Accordingly, for a given angular velocity, permanent magnet material, and air-gap height, *lg* one can say *E* ∝ *NphDl*, which means that a generator with a lower output voltage requires a lower value of *NphDl*. On the other hand, the minimum synchronous inductance, required for the boost CSI to operate in CCM, restricts the maximum number of poles, since *Ls* ∝ (*Dl*/*lg*)(*Nph*/*p*) 2 [33]. Hence, the desired values of the generator output voltage, (*E*<sup>2</sup> − (ω*sLsI*)<sup>2</sup> ) 1/2 ≅ *E* and synchronous inductance, *Ls*, are herein used as the PMSG design inputs for a given rated power.

## **3. Generator design and comparison**

In this section, a 1.5 MW PMSG is designed for the proposed DDWT topology and then the design is compared with an existing 1.5 MW PMSG [34].

#### **3.1 New generator design**

The generator is designed using GenAC toolbox of MagneForce Finite Element (FE) software. The designed machine is a low-speed, concentrated overlapping tooth coil double layered winding, three-phase PMSG rated at 1.5 MW. The generator is designed to be radial flux, cylindrical rotor, rotor surface mounted pole machine so that it can be compared with the existing 1.5 MW PMSG [34]. For the sake of comparison, both the existing and new 1.5 MW systems have the rated speed of 19.65 rpm, and the same generator airgap of 0.6 mm. Also, the line-to-line voltage of the new generator is 450 V, while the existing generator line-to-line voltage is 770 V, see **Table 1**.

**Figure 3** shows the cross-sectional view of the new generator, including the stator slot and tooth dimensions in details. Moreover, the insulation thickness in each slot is 3.5 mm. The overall slot-fill is about 62%. Also, the rotor shaft bore diameter is 700 mm. The stator bore diameter is 3500 mm, the stator length is 600 mm, and the airgap length, *lg*, is 0.6 mm. The PM material is sintered Neodymium Iron Boron (NdFeB) with the maximum residual flux density of *Br* = 1.2 T, and maximum coercive force of 24 k Oersted (i.e. *Hc* =1910 kA/m). The permanent magnet pole dimension is as 260 × 100mm<sup>2</sup> with a pole arc to pole pitch ratio of 0.8. **Figure 4** demonstrates the electromagnetic flux and the flux density distribution over the


**Table 1.**

*The existing and new generator input parameters.*

#### **Figure 3.**

*Cross sectional view of the new design 1.5 MW PMSG.*

PMSG at 1 *p*.*u*. load. As can be seen in **Figure 4**, the generator is not saturated at rated load. The maximum flux density is observed at the stator tooth, and the maximum airgap flux density measured using FE computations is about 1.15 T. Using circuit simulation for the new design, ~20 mH is needed to keep the dc-link current in CCM for different load levels.

#### **3.2 Generator parameter comparison**

In this subsection, the designed PMSG is compared with an existing 1.5 MW PM generator for DDWT [34]. **Table 2** shows the differences between the new and existing design parameters. In order to compare the two generators, the existing generator design is duplicated using the FE software. The design parameters for the existing generator are obtained from [34]. The comparison of the generator no-load line-to-line voltage, and the flux linkage with respect to rotor electrical position is presented in **Figure 5**. The no-load line-to-line voltage waveforms are shown in **Figure 5(a)**. The flux linkage of phase-A of the generator versus the rotor electrical position is presented in **Figure 5(b)**. As can be seen in **Figure 5(b)** that there is minimal distortion in the flux linkage, and it is sinusoidal for the generator designed. Furthermore, **Figure 5(b)** shows that the flux linkage is higher for the new design than the existing generator, as the new design has higher per phase inductance than that of the existing generator. It should be noted that even though the flux is higher in the new generator design, the maximum flux density is less than that permissible for the core, i.e. 1.2 T. As shown in **Table 2**, the number of

**49**

**Table 2.**

**Figure 4.**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

*Flux and flux density distribution (T) over the cross-section of the designed PMSG at full load.*

No. of poles, *p* 78 32 Airgap flux density, *Bmax* 1.01 T 1.15 T Winding current density, *J* 4 A/mm2 6 A/mm2 Linear current density 31.08 kA/m 36.60 kA/m Stator winding factor, *kw* 0.942 0.951 Stator q-axis inductance, *Lq* 12.20 mH 20.16 mH Stator d-axis inductance, *Ld* 12.19 mH 20.11 mH Stator inner diameter 4462 mm 3500 mm Stack length 500 mm 600 mm Number of slots 234 96 Winding turns per phase 56 30 Copper weight 384.4 kg 401.3 kg Core weight (stator) 21,552 kg 19,429 kg Rotor weight 40,990 kg 35,870 kg

**Parameter Existing generator Designed (new) generator**

poles in the existing generator is 78 which has been reduced to less than half in the

design, whereas, in the existing PMSG, each pole volume is 500 × 142 × 100 mm3

Thus, the total permanent magnet material is reduced by 9.8% in the new design. The reduction in the permanent magnet material reduces the dependency on the

The FE measured linear current density for the new design is 36.6 kA/m at the pole surface, as shown in **Table 2**, while the NdFeB coercive force is 1910 kA/m. In **Table 3**, the demagnetization of PM poles is computed for different output power factors and the results are presented. The computation of demagnetization is done using the MagneForce FE software. The software calculates the total flux with

in the new

.

new design to 32 poles. Also, each pole volume is 600 × 260 × 100 mm3

imported and highly unstable market of NdFeB materials.

*Comparison of the designed vs. existing generator.*

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

#### **Figure 4.**

*Rotating Machinery*

**Table 1.**

**48**

**Figure 3.**

*Cross sectional view of the new design 1.5 MW PMSG.*

*The existing and new generator input parameters.*

in CCM for different load levels.

**3.2 Generator parameter comparison**

PMSG at 1 *p*.*u*. load. As can be seen in **Figure 4**, the generator is not saturated at rated load. The maximum flux density is observed at the stator tooth, and the maximum airgap flux density measured using FE computations is about 1.15 T. Using circuit simulation for the new design, ~20 mH is needed to keep the dc-link current

**Parameter Existing generator Designed (new) generator**

Power rating 1.5 MW 1.5 MW Speed 19.65 rpm 19.65 rpm Air gap length 0.6 mm 0.6 mm Line-to-line voltage 770 V 450 V

In this subsection, the designed PMSG is compared with an existing 1.5 MW PM

generator for DDWT [34]. **Table 2** shows the differences between the new and existing design parameters. In order to compare the two generators, the existing generator design is duplicated using the FE software. The design parameters for the existing generator are obtained from [34]. The comparison of the generator no-load line-to-line voltage, and the flux linkage with respect to rotor electrical position is presented in **Figure 5**. The no-load line-to-line voltage waveforms are shown in **Figure 5(a)**. The flux linkage of phase-A of the generator versus the rotor electrical position is presented in **Figure 5(b)**. As can be seen in **Figure 5(b)** that there is minimal distortion in the flux linkage, and it is sinusoidal for the generator designed. Furthermore, **Figure 5(b)** shows that the flux linkage is higher for the new design than the existing generator, as the new design has higher per phase inductance than that of the existing generator. It should be noted that even though the flux is higher in the new generator design, the maximum flux density is less than that permissible for the core, i.e. 1.2 T. As shown in **Table 2**, the number of

*Flux and flux density distribution (T) over the cross-section of the designed PMSG at full load.*


#### **Table 2.**

*Comparison of the designed vs. existing generator.*

poles in the existing generator is 78 which has been reduced to less than half in the new design to 32 poles. Also, each pole volume is 600 × 260 × 100 mm3 in the new design, whereas, in the existing PMSG, each pole volume is 500 × 142 × 100 mm3 . Thus, the total permanent magnet material is reduced by 9.8% in the new design. The reduction in the permanent magnet material reduces the dependency on the imported and highly unstable market of NdFeB materials.

The FE measured linear current density for the new design is 36.6 kA/m at the pole surface, as shown in **Table 2**, while the NdFeB coercive force is 1910 kA/m. In **Table 3**, the demagnetization of PM poles is computed for different output power factors and the results are presented. The computation of demagnetization is done using the MagneForce FE software. The software calculates the total flux with

**Figure 5.**

*Generator (a) line-to-line voltage, and (b) phase flux linkage vs. rotor electrical position obtained through FE computations for the existing and new design of the generators.*


#### **Table 3.**

*Percent demagnetization of pm poles for varying output power factor and s = 1 p.u. using FE computations.*

virgin (non-demagnetized) magnets. It then runs the simulation and keeps track of the worst demagnetization that each element within the rotor pole experiences throughout a complete ac cycle. The FE software uses the demagnetized elements to re-construct the magnet, and then, calculates the total flux from this demagnetized magnet. Finally, the software calculates the percentage drop in total flux from the virgin magnets to the demagnetized magnets and reports the difference as the amount of demagnetization. It should be noted that the demagnetization analysis is based on loading of the generators and total flux drop, and it does not take into account the thermal models for the generators.

A normalized comparison of the copper and core losses, rotor and stator weights, and the volume of the two generators is presented in **Figure 6**. In **Table 4**, the efficiency, as well as core and copper losses, are provided for different power factors and a constant apparent power, i.e. *S* =1 p.u. The base values for the per-unit (p.u.) calculations are *Sbase* =1.5 MVA, and *Vbase* =690 V. As can be observed, the copper loss in the new generator increases. This is due to the higher current level in the low-voltage generator. Although the new design has lower number of turns in the generator winding – lower overall stator resistance – the line current is higher, which results in higher copper loss in the new low voltage generator. Nevertheless, the core loss for the designed low voltage generator decreases by about 9.5%. This reduction in the core loss in the new designed generator is due to the reduction in the number of poles in the new designed generator, which results in a substantial decrease in the generator output voltage and frequency, see **Table 2**. It can be observed that the generator efficiency is not compromised with full load efficiency of the new and existing generator being 94.4 and 93.8%, respectively, while the stator and rotor dimensions are significantly decreased. The no-load core-losses for the existing and new generator designs are computed to be 0.040 p.u. and 0.0351 p.u., respectively. The change in core losses from no-load to full load can be attributed to the armature reaction, see **Table 4**. It should be noted that the losses and efficiency data presented in in **Table 4**, and **Figure 6** are

**51**

in the following sections.

**4. System controls**

stabilize the input dc-voltage.

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

*Comparison of normalized (a) copper loss, (b) core loss, (c) stator diameter, (d) rotor weight, (e) stator weight, and (f) volume of the existing and designed PM generator for DDWT using FE computation.*

Copper Loss (*p*.*u*.) Existing 0.0245 0.0245 0.0250

Core Loss (*p*.*u*.) Existing 0.0405 0.0405 0.0405

Efficiency (%) Existing 93.50 93.49 93.45

*pf* **= 0.95 0.85 0.75**

New 0.0268 0.0269 0.0278

New 0.0353 0.0354 0.0354

New 93.79 93.77 93.68

obtained using the FE models of the PMSGs. A comparison of the rotor and stator weights, respectively, for the two generators is presented in **Figure 6**. The rotor weight is the cumulative weight of the NdFeB PM poles and the rotor core. Similarly, the stator weight presented in the combined weight of the armature windings and the stator core. The stator and rotor weight reductions for the new generator are ~9.4 and 12.5%, respectively. Additionally, the overall volume of the new generator is reduced by ~ 2.79% than the existing design. It can be observed that the decrease in the stator diameter (which is about 10%) does not render to a comparable reduction in the overall volume of the new generator. This is due to the higher stack length of the new generator. The designed generator has a stack length to diameter (*l*/*D*) ratio of 0.13 as compared to the existing generator design with *l*/*D* ratio of 0.11. The decrease in the generator volume and weight along with the elimination the failure prone dc-bus capacitor and circumventing the dc-link inductor needed in CSIs will further render into decreased capital cost of the overall system. The simulation and experimental results evaluating and validating the performance of the proposed topology of the DDWT are presented

*Core loss, copper loss, and efficiency for varying output power factor and S = 1 p.u. using FE computations.*

In this section, first, the control technique implemented for the developed topology is briefly explained. The developed system is controlled with generator side VSC being modulated for power transfer to the grid and the boost inverter modulated to

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

**Figure 6.**

**Table 4.**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

#### **Figure 6.**

*Rotating Machinery*

**Figure 5.**

**Table 3.**

V

L L (V)

Flux (VSec.)


> -50 0 50

*computations for the existing and new design of the generators.*

virgin (non-demagnetized) magnets. It then runs the simulation and keeps track of the worst demagnetization that each element within the rotor pole experiences throughout a complete ac cycle. The FE software uses the demagnetized elements to re-construct the magnet, and then, calculates the total flux from this demagnetized magnet. Finally, the software calculates the percentage drop in total flux from the virgin magnets to the demagnetized magnets and reports the difference as the amount of demagnetization. It should be noted that the demagnetization analysis is based on loading of the generators and total flux drop, and it does not take into

*Percent demagnetization of pm poles for varying output power factor and s = 1 p.u. using FE computations.*

(a) 0 180 360 540 720

Existing design New design

*pf* **= 0.95 0.85 0.75 Demagnetization (%)**

(b) Rotor position (el. Deg.)

*Generator (a) line-to-line voltage, and (b) phase flux linkage vs. rotor electrical position obtained through FE* 

Existing design 0.080 0.075 0.078 New design 0.084 0.080 0.081

0 180 360 540 720

A normalized comparison of the copper and core losses, rotor and stator weights, and the volume of the two generators is presented in **Figure 6**. In **Table 4**, the efficiency, as well as core and copper losses, are provided for different power factors and a constant apparent power, i.e. *S* =1 p.u. The base values for the per-unit (p.u.) calculations are *Sbase* =1.5 MVA, and *Vbase* =690 V. As can be observed, the copper loss in the new generator increases. This is due to the higher current level in the low-voltage generator. Although the new design has lower number of turns in the generator winding – lower overall stator resistance – the line current is higher, which results in higher copper loss in the new low voltage generator. Nevertheless, the core loss for the designed low voltage generator decreases by about 9.5%. This reduction in the core loss in the new designed generator is due to the reduction in the number of poles in the new designed generator, which results in a substantial decrease in the generator output voltage and frequency, see **Table 2**. It can be observed that the generator efficiency is not compromised with full load efficiency of the new and existing generator being 94.4 and 93.8%, respectively, while the stator and rotor dimensions are significantly decreased. The no-load core-losses for the existing and new generator designs are computed to be 0.040 p.u. and 0.0351 p.u., respectively. The change in core losses from no-load to full load can be attributed to the armature reaction, see **Table 4**. It should be noted that the losses and efficiency data presented in in **Table 4**, and **Figure 6** are

account the thermal models for the generators.

**50**

*Comparison of normalized (a) copper loss, (b) core loss, (c) stator diameter, (d) rotor weight, (e) stator weight, and (f) volume of the existing and designed PM generator for DDWT using FE computation.*


#### **Table 4.**

*Core loss, copper loss, and efficiency for varying output power factor and S = 1 p.u. using FE computations.*

obtained using the FE models of the PMSGs. A comparison of the rotor and stator weights, respectively, for the two generators is presented in **Figure 6**. The rotor weight is the cumulative weight of the NdFeB PM poles and the rotor core. Similarly, the stator weight presented in the combined weight of the armature windings and the stator core. The stator and rotor weight reductions for the new generator are ~9.4 and 12.5%, respectively. Additionally, the overall volume of the new generator is reduced by ~ 2.79% than the existing design. It can be observed that the decrease in the stator diameter (which is about 10%) does not render to a comparable reduction in the overall volume of the new generator. This is due to the higher stack length of the new generator. The designed generator has a stack length to diameter (*l*/*D*) ratio of 0.13 as compared to the existing generator design with *l*/*D* ratio of 0.11. The decrease in the generator volume and weight along with the elimination the failure prone dc-bus capacitor and circumventing the dc-link inductor needed in CSIs will further render into decreased capital cost of the overall system. The simulation and experimental results evaluating and validating the performance of the proposed topology of the DDWT are presented in the following sections.

### **4. System controls**

In this section, first, the control technique implemented for the developed topology is briefly explained. The developed system is controlled with generator side VSC being modulated for power transfer to the grid and the boost inverter modulated to stabilize the input dc-voltage.

**Figure 7.** *Block diagram of controller for the VSC [26].*

#### **Figure 8.**

*Block diagram of controller for the single-stage boost inverter [26].*

The power electronics interface proposed for the DDWT drivetrain is controlled to regulate the power injected to the grid and to maintain the average dc-side voltage. The generator side VSC is modulated to inject desired power to the dc-side.

**53**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

**Figure 7** shows the block diagram of the controller for VSC. The generator line-toline voltages and dc-side voltage and current are the measured feedback received by this controller. The ac side signals are then converted to the *dqo* reference frame.

on the dc side. The error is then mitigated by using a Proportional-Integral (PI)

generated *q*- axis voltage and the measured *d*-axis voltages serve as base to generate the switching signal for the VSC. The controller is equipped with one of the existing

<sup>3</sup> , [36, 37]. **Figure 8** shows the block diagram of the controller for the boost CSI. The controls for this converter help regulate the dc-side voltage and to modulate the reactive power injected into the grid. Thus, the controller for the VSC helps modulate the active power injected into the grid and the controller for the boost CSI modulates the reactive power into the grid. The inputs to the boost CSI controller are the measured grid-side line-to-line voltages, line currents and dc-side. The three-phase signals are converted to the *dqo* reference frame. The *dqo* quantities are used to compute the reactive power injected to the grid, which is compared to the reference reactive power. The error is then mitigated using a PI controller which generates the reference angle, θ for the boost inverter. Similarly, the dc-side voltage, *V dc* (where,*V dc* = 〈*v dc*〉) is compared to the desired dc-side voltage, *V dc* <sup>∗</sup> and a PI controller is used to generate the modulation index, *D* for the implementing the PPWM switching technique [26]. The details on the PPWM technique has been provided in [29–31].

This section presents the simulation results for the evaluation of the feasibility and performance of the developed topology of the DDWT. The software platform used for these evaluations is MATLAB/Simulink environment using SimPowerSystems toolbox. The simulated system uses the FE design of the generator as presented in Section 3. The switching frequency of 6 kHz is used for both the VSC and the boost CSI. The switching methodology for the VSC is the Sine PWM (SPWM) technique while the boost CSI employs the PPWM switching technique [29, 30]. As described in the previous section, the VSC controller modulates the active power transfer injected into the grid and the boost CSI controller regulates the dc-side voltage and controls the reactive power injected into the grid. A detailed discussion on these controllers and their performance is presented in [26]. The

output filter parameters for the boost CSI are C*<sup>f</sup>* = 20 μF and *Lf* = 5 mH.

The quality of the output waveforms of the proposed system is demonstrated in **Figures 9** and **10**. **Figure 10** presents the steady-state waveforms of generator lineto-line voltage and line current when the generator speed is 18.75 rpm. The power converters were modulated to inject 1 MW (0.67 p.u.) P into the grid at unity power factor (UPF). The boost CSI controller was regulating the dc-side voltage at 500 V. **Figure 10(a)** and **(b)** show that the THD of the generator output line current is ~1.2%, which further ensures less harmonics transferred to the generator side and eliminates the low-frequency torque ripple on the generator shaft. **Figure 9** presents the waveforms of the boost CSI output line-to-line voltage at the point of common coupling, and line current injected into the grid. The grid line current THD is computed to be ~3.3%, which is complaint of the IEEE 1547–2018 interconnection standards [38]. **Figure 11** presents the dynamic behavior of the proposed DDWT system. The generator speed, three-phase line-to-line voltage waveforms at the point of common coupling (PCC), three-phase generator line current waveforms, and the dc-side current are shown in **Figure 11**. The first event occurs at *t* = 1*s*, when the generator speed

controller which then generates the desired *q*-axis converter voltage, *V <sup>q</sup>*

is compared to the measured power

*VSC*∗

. The

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

The desired active power on the grid side, *P* <sup>∗</sup>

MPPT techniques, e.g. *P* <sup>∗</sup> = *K opt*ω *<sup>R</sup>*

**5. Simulation results**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

**Figure 7** shows the block diagram of the controller for VSC. The generator line-toline voltages and dc-side voltage and current are the measured feedback received by this controller. The ac side signals are then converted to the *dqo* reference frame. The desired active power on the grid side, *P* <sup>∗</sup> is compared to the measured power on the dc side. The error is then mitigated by using a Proportional-Integral (PI) controller which then generates the desired *q*-axis converter voltage, *V <sup>q</sup> VSC*∗ . The generated *q*- axis voltage and the measured *d*-axis voltages serve as base to generate the switching signal for the VSC. The controller is equipped with one of the existing MPPT techniques, e.g. *P* <sup>∗</sup> = *K opt*ω *<sup>R</sup>* <sup>3</sup> , [36, 37].

**Figure 8** shows the block diagram of the controller for the boost CSI. The controls for this converter help regulate the dc-side voltage and to modulate the reactive power injected into the grid. Thus, the controller for the VSC helps modulate the active power injected into the grid and the controller for the boost CSI modulates the reactive power into the grid. The inputs to the boost CSI controller are the measured grid-side line-to-line voltages, line currents and dc-side. The three-phase signals are converted to the *dqo* reference frame. The *dqo* quantities are used to compute the reactive power injected to the grid, which is compared to the reference reactive power. The error is then mitigated using a PI controller which generates the reference angle, θ for the boost inverter. Similarly, the dc-side voltage, *V dc* (where,*V dc* = 〈*v dc*〉) is compared to the desired dc-side voltage, *V dc* <sup>∗</sup> and a PI controller is used to generate the modulation index, *D* for the implementing the PPWM switching technique [26]. The details on the PPWM technique has been provided in [29–31].

## **5. Simulation results**

*Rotating Machinery*

**Figure 7.**

*Block diagram of controller for the VSC [26].*

**52**

**Figure 8.**

*Block diagram of controller for the single-stage boost inverter [26].*

The power electronics interface proposed for the DDWT drivetrain is controlled

to regulate the power injected to the grid and to maintain the average dc-side voltage. The generator side VSC is modulated to inject desired power to the dc-side.

This section presents the simulation results for the evaluation of the feasibility and performance of the developed topology of the DDWT. The software platform used for these evaluations is MATLAB/Simulink environment using SimPowerSystems toolbox. The simulated system uses the FE design of the generator as presented in Section 3. The switching frequency of 6 kHz is used for both the VSC and the boost CSI. The switching methodology for the VSC is the Sine PWM (SPWM) technique while the boost CSI employs the PPWM switching technique [29, 30]. As described in the previous section, the VSC controller modulates the active power transfer injected into the grid and the boost CSI controller regulates the dc-side voltage and controls the reactive power injected into the grid. A detailed discussion on these controllers and their performance is presented in [26]. The output filter parameters for the boost CSI are C*<sup>f</sup>* = 20 μF and *Lf* = 5 mH.

The quality of the output waveforms of the proposed system is demonstrated in **Figures 9** and **10**. **Figure 10** presents the steady-state waveforms of generator lineto-line voltage and line current when the generator speed is 18.75 rpm. The power converters were modulated to inject 1 MW (0.67 p.u.) P into the grid at unity power factor (UPF). The boost CSI controller was regulating the dc-side voltage at 500 V. **Figure 10(a)** and **(b)** show that the THD of the generator output line current is ~1.2%, which further ensures less harmonics transferred to the generator side and eliminates the low-frequency torque ripple on the generator shaft. **Figure 9** presents the waveforms of the boost CSI output line-to-line voltage at the point of common coupling, and line current injected into the grid. The grid line current THD is computed to be ~3.3%, which is complaint of the IEEE 1547–2018 interconnection standards [38].

**Figure 11** presents the dynamic behavior of the proposed DDWT system. The generator speed, three-phase line-to-line voltage waveforms at the point of common coupling (PCC), three-phase generator line current waveforms, and the dc-side current are shown in **Figure 11**. The first event occurs at *t* = 1*s*, when the generator speed

#### **Figure 9.**

*Simulated results of inverter output (a) line current waveform, (b) line current FFT analysis, (c) lineto-line voltage waveform, and (d) line-to-line voltage FFT analysis for grid-tied system the generator speed is 18.75 rpm and the system is injecting 0.67p.u. Active power into the grid [35].*

#### **Figure 10.**

*Simulated results of generator output (a) line current waveform, (b) line current FFT analysis, (c) line-to-line voltage waveform, and (c) line-to-line voltage FFT analysis, for grid-tied system when the generator speed is 18.75 rpm and it is supplying [35].*

**55**

**Figure 12.**

*Laboratory scale 1.5 kW, setup for the proposed DDWT topology [35].*

**Figure 11.**

*the grid [35].*

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

*Plot of (a)PMSG speed, (b) three-phase line-to-line voltage at the PCC, (c) three-phase generator line current, and (d) dc-side current showing system behavior for varying wind speed and three-phase low-voltage fault at* 

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

#### **Figure 11.**

*Rotating Machinery*

1.5

I

L o u t (p.u.)

V

**Figure 9.**

I

V

L L g e n (p.u.)


*18.75 rpm and it is supplying [35].*

0

1

L g e n (p.u.)


0

2

L L o u t (p.u.)


0

2


0

(a) 0 20 40 60 80 100

> (c) Time (ms)

*18.75 rpm and the system is injecting 0.67p.u. Active power into the grid [35].*

(a) 0. 0.2 0.4 0.6 0.8 1

> (c) Time (s)

0 0.2 0.4 0.6 0.8 1

0 20 40 60 80 100



> | FFT (I L g e n ) | (%)

> | FFT (V L L g e n ) | (%)

*Simulated results of generator output (a) line current waveform, (b) line current FFT analysis, (c) line-to-line voltage waveform, and (c) line-to-line voltage FFT analysis, for grid-tied system when the generator speed is* 

0

2

4

0

2

4

*Simulated results of inverter output (a) line current waveform, (b) line current FFT analysis, (c) lineto-line voltage waveform, and (d) line-to-line voltage FFT analysis for grid-tied system the generator speed is* 

0

2.5

0

5

2.5

5

(b) 200 400 600 800 1000

> (d) Frequency (Hz)

(b) 200 400 600 800 1000

> (d) Frequency (Hz)

200 400 600 800 1000

200 400 600 800 1000

**54**

**Figure 10.**

*Plot of (a)PMSG speed, (b) three-phase line-to-line voltage at the PCC, (c) three-phase generator line current, and (d) dc-side current showing system behavior for varying wind speed and three-phase low-voltage fault at the grid [35].*

changes gradually from 7.5 rpm to 19.65 rpm. The sudden speed changes shown in **Figure 11** are not realistic and is just used as a case to evaluate the robustness of the system controls. **Figure 11(c)** and **(d)** show that the system easily tracks the maximum available wind power. The fault ride-through capability of the system is also evaluated. This is done by simulating a three-phase low-voltage fault on the inverter terminals at *t* = 2 *s*, where the grid voltage decreases by 35% and the fault is cleared after 0.5*s* [39]. The developed system rides-through this low-voltage seamlessly and starts normal operation again in ~0.38s after the clearing of the fault.

## **6. Experimental evaluation results**

The feasibility of the developed DDWT topology is evaluated using experimental results from 1.5 kW, 240VLL laboratory scale prototype, in this section (see **Figure 12**). In this setup, a motor-drive system is used as the wind turbine emulator that runs a commercially available PMSG, which feeds a three-phase 50Ω resistive load through the proposed back-to-back converter and a CL filter. The filter values are *Cf* =20 *μ*F and *Lf* =5 mH, when *fs* <sup>=</sup> 6 kHz for the boost CSI. For the first test case to evaluate if the generator synchronous inductance can be utilized in place of the dc-link inductor, a diode bridge rectifier is used as the generator side. The waveforms of the generator line current, generator line-to-line voltage, load line current, and load line-to-line voltage are shown in **Figure 13**. The PMSG produces rms line-to-line voltage of 45 V at 30 Hz. This is boosted to an output load line-to-line voltage of 208 V at 60 Hz by the boost CSI. The generator current and load current

#### **Figure 13.**

*Experimentally obtained waveforms of the (a) generator line current, (b) generator line-to-line voltage, (c) load line current, and (d) load line-to-line voltage when the generator is rotating at 450 rpm (30 Hz), the output is regulated at 208 V, and the generator side converter is a diode bridge rectifier [35].*

**57**

the system output.

**Figure 14.**

**7. Conclusion**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

THDs are measured to be 33.1 and 3.0%, respectively. For the next test case, the generator side converter employed is a VSC. **Figure 14** shows the waveforms of the same quantities for this test case. The switching frequency for the VSC is *fs* =25 kHz. The average dc side voltage was maintained at 70 V. The generator current and load current THDs are measured to be 2.9 and 2.8%, respectively The evaluation results presented in **Figures 13** and **14** demonstrate the capability of the developed systems to utilize the generator synchronous inductance, while still maintain the quality of

*Experimentally obtained waveforms of the (a) generator line current, (b) generator line-to-line voltage, (c) load line current, and (d) load line-to-line voltage when the generator is rotating at 450 rpm (30 Hz) and* 

*the output is regulated at 208 V with the generator side converter as VSC [35].*

In this chapter, an inductorless generator-converter topology has been presented for DDWTs. This paper also presents a new topology of power electronics interface for the DDWTs. The grid side VSC used in a conventional DDWT has been replaced by a boost CSI. The dc-link inductor inherently required in a CSI topology has been eliminated by using the synchronous inductance of the PMSG. This topology has further facilitated the elimination of vulnerable dc-bus electrolytic capacitors, thereby increasing the overall system reliability. Furthermore, less frequent failures of the capacitors will lead to increase in system availability, resulting in substantial decrease in the maintenance costs. In this chapter, FE analysis is used to design and evaluated a 1.5 MW generator for the proposed system. The FE analysis is also used to compare the designed and an actual generator. The advantages of using the proposed topology have been demonstrated through reduction in generator weight,

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

**Figure 14.**

*Rotating Machinery*

**56**

**Figure 13.**

*Experimentally obtained waveforms of the (a) generator line current, (b) generator line-to-line voltage, (c) load line current, and (d) load line-to-line voltage when the generator is rotating at 450 rpm (30 Hz), the* 

changes gradually from 7.5 rpm to 19.65 rpm. The sudden speed changes shown in **Figure 11** are not realistic and is just used as a case to evaluate the robustness of the system controls. **Figure 11(c)** and **(d)** show that the system easily tracks the maximum available wind power. The fault ride-through capability of the system is also evaluated. This is done by simulating a three-phase low-voltage fault on the inverter terminals at *t* = 2 *s*, where the grid voltage decreases by 35% and the fault is cleared after 0.5*s* [39]. The developed system rides-through this low-voltage seamlessly and

The feasibility of the developed DDWT topology is evaluated using experimental results from 1.5 kW, 240VLL laboratory scale prototype, in this section (see **Figure 12**). In this setup, a motor-drive system is used as the wind turbine emulator that runs a commercially available PMSG, which feeds a three-phase 50Ω resistive load through the proposed back-to-back converter and a CL filter. The filter values are *Cf* =20 *μ*F and *Lf* =5 mH, when *fs* <sup>=</sup> 6 kHz for the boost CSI. For the first test case to evaluate if the generator synchronous inductance can be utilized in place of the dc-link inductor, a diode bridge rectifier is used as the generator side. The waveforms of the generator line current, generator line-to-line voltage, load line current, and load line-to-line voltage are shown in **Figure 13**. The PMSG produces rms line-to-line voltage of 45 V at 30 Hz. This is boosted to an output load line-to-line voltage of 208 V at 60 Hz by the boost CSI. The generator current and load current

starts normal operation again in ~0.38s after the clearing of the fault.

**6. Experimental evaluation results**

*output is regulated at 208 V, and the generator side converter is a diode bridge rectifier [35].*

*Experimentally obtained waveforms of the (a) generator line current, (b) generator line-to-line voltage, (c) load line current, and (d) load line-to-line voltage when the generator is rotating at 450 rpm (30 Hz) and the output is regulated at 208 V with the generator side converter as VSC [35].*

THDs are measured to be 33.1 and 3.0%, respectively. For the next test case, the generator side converter employed is a VSC. **Figure 14** shows the waveforms of the same quantities for this test case. The switching frequency for the VSC is *fs* =25 kHz. The average dc side voltage was maintained at 70 V. The generator current and load current THDs are measured to be 2.9 and 2.8%, respectively The evaluation results presented in **Figures 13** and **14** demonstrate the capability of the developed systems to utilize the generator synchronous inductance, while still maintain the quality of the system output.

### **7. Conclusion**

In this chapter, an inductorless generator-converter topology has been presented for DDWTs. This paper also presents a new topology of power electronics interface for the DDWTs. The grid side VSC used in a conventional DDWT has been replaced by a boost CSI. The dc-link inductor inherently required in a CSI topology has been eliminated by using the synchronous inductance of the PMSG. This topology has further facilitated the elimination of vulnerable dc-bus electrolytic capacitors, thereby increasing the overall system reliability. Furthermore, less frequent failures of the capacitors will lead to increase in system availability, resulting in substantial decrease in the maintenance costs. In this chapter, FE analysis is used to design and evaluated a 1.5 MW generator for the proposed system. The FE analysis is also used to compare the designed and an actual generator. The advantages of using the proposed topology have been demonstrated through reduction in generator weight,

volume, and amount of PM material. The controls for the proposed system have also been discussed in this chapter. The feasibility and performance of the proposed DDWT system has been validated through simulation results and experimental results from a laboratory scale prototype, in this chapter.

## **Author details**

Akanksha Singh National Renewable Energy Laboratory, Golden, USA

\*Address all correspondence to: akanksha.singh@nrel.gov

© 2019 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

**59**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

Challenges, design tools, and opportunities. IEEE Industrial Electronics Magazine. 2013;**7**(2):17-26

2016. pp. 1-4

2012

[11] Chen J, Chen W, Li J, Zhang X, Sun P. Lifetime assessment of DC link electrolytic capacitor of wind power converter based on operational condition. In: 2016 IEEE International Conference on High Voltage Engineering and Application (ICHVE), Chengdu.

[12] Leban K, Ritchie E, Argeseanu A. Design preliminaries for direct drive under water wind turbine generator. In: IEEE International Conference on Electrical Machines. Marseille, France;

[13] Mcmillan D, Ault GW. Technoeconomic comparison of operational

gearbox-driven wind turbines. IEEE Transactions on Energy Conversion.

[14] Chen Z, Guerrero JM, Blaabjerg F. A review of the state of the art of power electronics for wind turbines. IEEE Transactions on Power Electronics.

aspects for direct drive and

2010;**25**(1):191-198

2009;**24**(8):1859-1875

2018;**2018**(1, 1):10-16

[15] Singh A, Mirafzal B. Indirect boost matrix converter and lowvoltage generator for direct drive wind turbines. The Journal of Engineering.

[16] Singh A, Mirafzal B. A generatorconverter design for direct drive wind turbines. In: Proc. IEEE ECCE. 2016

[17] Blaabjerg F, Liserre M, Ma K. Power electronics for wind turbine systems. IEEE Transactions on Industry Applications. 2012;**48**(2):708-719

[18] Lumbreras C, Guerrero JM, Garcia P, Briz F, Reigosa DD. Control of a small wind turbine in the high wind speed

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

[1] Global Wind Energy Council Report. [Online] Available from: http://www.gwec.net/global-figures/

from: http://www.awea.org

United Kingdom: Wiley; 2011

[4] Ribrant J, Bertling L. Survey of failures in wind power systems with focus on swedish wind power plants during 1997-2005. IEEE Transactions on Energy Conversion. 2007;**22**(1):167-173

[5] Daneshi-Far Z, Capolino GA, Henao H. Review of failures and condition monitoring in wind turbine generators.

[6] Tavner P, Bussel G, Spinato F. Machine and converter reliabilities in wind turbines. In: IEEE International Conference on Power Electronics, Machines and Drives.

[7] Polinder H, Van Der Pijl FFA, de Vilder GJ, Tavner P. Comparison of direct-drive and geared generator concepts for wind turbines. IEEE Transactions on Energy Conversion. 2006;**21**(3):725-733

[8] Spinato F, Tavner PJ, van Bussel GJW, Koutoulakos E. Reliability of wind turbine subassemblies. IET Renewable Power Generation. 2009;**3**(4):387-401

[9] Liserre M, Cardenas R, Molinas M, Rodriguez J. Overview of multi-MW wind turbines and wind parks. IEEE Transactions on Industrial Electronics.

[10] Wang H, Liserre M, Blaabjerg F. Toward reliable power electronics:

In: International Conference on Electrical Machines. 2010

Dublin, Ireland; 2006

2011;**58**(4):1081-1095

[2] AWEA U.S. Wind Industry Market Reports for 2015 [Online]. Available

[3] Manwell J, Mcgowan J, Rogers A. Wind Energy Explained: Theory, Design, and Applications. Second ed.

**References**

wind-in-numbers/

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

## **References**

*Rotating Machinery*

**58**

**Author details**

Akanksha Singh

provided the original work is properly cited.

National Renewable Energy Laboratory, Golden, USA

\*Address all correspondence to: akanksha.singh@nrel.gov

© 2019 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium,

volume, and amount of PM material. The controls for the proposed system have also been discussed in this chapter. The feasibility and performance of the proposed DDWT system has been validated through simulation results and experimental

results from a laboratory scale prototype, in this chapter.

[1] Global Wind Energy Council Report. [Online] Available from: http://www.gwec.net/global-figures/ wind-in-numbers/

[2] AWEA U.S. Wind Industry Market Reports for 2015 [Online]. Available from: http://www.awea.org

[3] Manwell J, Mcgowan J, Rogers A. Wind Energy Explained: Theory, Design, and Applications. Second ed. United Kingdom: Wiley; 2011

[4] Ribrant J, Bertling L. Survey of failures in wind power systems with focus on swedish wind power plants during 1997-2005. IEEE Transactions on Energy Conversion. 2007;**22**(1):167-173

[5] Daneshi-Far Z, Capolino GA, Henao H. Review of failures and condition monitoring in wind turbine generators. In: International Conference on Electrical Machines. 2010

[6] Tavner P, Bussel G, Spinato F. Machine and converter reliabilities in wind turbines. In: IEEE International Conference on Power Electronics, Machines and Drives. Dublin, Ireland; 2006

[7] Polinder H, Van Der Pijl FFA, de Vilder GJ, Tavner P. Comparison of direct-drive and geared generator concepts for wind turbines. IEEE Transactions on Energy Conversion. 2006;**21**(3):725-733

[8] Spinato F, Tavner PJ, van Bussel GJW, Koutoulakos E. Reliability of wind turbine subassemblies. IET Renewable Power Generation. 2009;**3**(4):387-401

[9] Liserre M, Cardenas R, Molinas M, Rodriguez J. Overview of multi-MW wind turbines and wind parks. IEEE Transactions on Industrial Electronics. 2011;**58**(4):1081-1095

[10] Wang H, Liserre M, Blaabjerg F. Toward reliable power electronics:

Challenges, design tools, and opportunities. IEEE Industrial Electronics Magazine. 2013;**7**(2):17-26

[11] Chen J, Chen W, Li J, Zhang X, Sun P. Lifetime assessment of DC link electrolytic capacitor of wind power converter based on operational condition. In: 2016 IEEE International Conference on High Voltage Engineering and Application (ICHVE), Chengdu. 2016. pp. 1-4

[12] Leban K, Ritchie E, Argeseanu A. Design preliminaries for direct drive under water wind turbine generator. In: IEEE International Conference on Electrical Machines. Marseille, France; 2012

[13] Mcmillan D, Ault GW. Technoeconomic comparison of operational aspects for direct drive and gearbox-driven wind turbines. IEEE Transactions on Energy Conversion. 2010;**25**(1):191-198

[14] Chen Z, Guerrero JM, Blaabjerg F. A review of the state of the art of power electronics for wind turbines. IEEE Transactions on Power Electronics. 2009;**24**(8):1859-1875

[15] Singh A, Mirafzal B. Indirect boost matrix converter and lowvoltage generator for direct drive wind turbines. The Journal of Engineering. 2018;**2018**(1, 1):10-16

[16] Singh A, Mirafzal B. A generatorconverter design for direct drive wind turbines. In: Proc. IEEE ECCE. 2016

[17] Blaabjerg F, Liserre M, Ma K. Power electronics for wind turbine systems. IEEE Transactions on Industry Applications. 2012;**48**(2):708-719

[18] Lumbreras C, Guerrero JM, Garcia P, Briz F, Reigosa DD. Control of a small wind turbine in the high wind speed

region. IEEE Transactions on Power Electronics. 2016;**30**(10):6980-6991

[19] Dai J, Xu DD, Wu B. A novel control scheme for current-source converterbased PMSG wind energy conversion systems. IEEE Transactions on Power Electronics. 2009;**24**(4):963-972

[20] Tenca P, Rockhill AA, Lipo TA, Tricoli P. Current source topology for wind turbines with decreased mains current harmonics, further reducible via functional minimization. IEEE Transactions on Power Electronics. 2008;**23**(3):1143-1155

[21] Lee JS, Lee KB, Blaabjerg F. Openswitch fault detection method of a back-to-back converter using NPC topology for wind turbine systems. IEEE Transactions on Industry Applications. 2015;**51**(1):325-335

[22] Blaabjerg F, Ma K. Future on power electronics for wind turbine systems. IEEE Journal of Emerging and Selected Topics in Power Electron. 2013;**1**(3):139-151

[23] Zavvos A, Mcdonald A, Mueller M. Optimisation tools for large permanent magnet generators for direct drive wind turbines. IET Renewable Power Generation. 2013;**7**(2):163-171

[24] Zhang Z, Zhao Y, Qiao W, Qu L. A space-vector-modulated sensorless direct-torque control for directdrive PMSG wind turbines. IEEE Transactions on Industry Applications. 2014;**50**(4):2331-2441

[25] Potgieter JHJ, Kamper MJ. Design optimization of directly grid-connected PM machines for wind energy applications. IEEE Transactions on Industry Applications. 2015;**51**(4):2949-2958

[26] Singh A, Mirafzal B. Three-phase single-stage boost inverter for direct drive wind turbines. In: 2016 IEEE

Energy Conversion Congress and Exposition (ECCE). Milwaukee, WI; 2016. pp. 1-7

[27] Singh A, Mirafzal B. A low-voltage generator-converter topology for direct drive wind turbines. In: 2016 IEEE 7th International Symposium on Power Electron. For Distributed Gener. Sys. (PEDG). Vancouver, BC; 2016. pp. 1-6

[28] Wang Z, Wu B, Xu D, Zargari NR. A current-source-converter-based high-power high-speed PMSM drive with 420-Hz switching frequency. IEEE Transactions on Industrial Electronics. 2012;**59**(7):2970-2981

[29] Mirafzal B, Saghaleini M, Kaviani A. An SVPWM-based switching pattern for stand-alone and grid-connected three-phase single-stage boost inverters. IEEE Transactions on Power Electronics. 2011;**26**(4):1102-1111

[30] Singh A, Milani AA, Mirafzal B. Modified phasor pulse width modulation method for three-phase single-stage boost inverter. In: 2014 IEEE Applied Power Electron. Conf. And Expo. - APEC 2014. Fort Worth, TX; 2014. pp. 1276-1280

[31] Singh A, Kaviani AK, Mirafzal B. On dynamics models and stability of three-phase phasor PWM based CSI for stand-alone applications. IEEE Transactions on Industrial Electronics. May 2015;**62**(5):2698-2707

[32] Holmes DG, Lipo TA. Pulse Width Modulation for Power Converters: Principles and Practice. First ed. New York, USA: John Wiley & Sons; 2003

[33] Fitzgerald AE, Kingsley C Jr, Umans SD. Electric Machinery. Sixth ed. New York, USA: McGraw Hill; 2003

[34] Northern Power NW 1500 Direct-Drive Generator –Report, [Online] Available from: http://www.nrel.gov/ docs/fy08osti/40177.pdf

**61**

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines*

*DOI: http://dx.doi.org/10.5772/intechopen.85877*

[35] Singh A, Benzaquen J, Mirafzal B. Current source generator–converter topology for direct-drive wind turbines. IEEE Transactions on Industry Applications. March-April

[36] Fateh F, White WN, Gruenbacher D. A maximum power tracking technique for grid-connected DFIGbased wind turbines. IEEE Journal of Emerging and Selected Topics in Power

Electronics. 2015;**3**(4):957-966

[37] Yuan X, Li Y. Control of variable pitch and variable speed direct-drive wind turbines in weak grid systems with active power balance. IET Renewable Power Generation. 2014;**8**(2):119-131

[38] IEEE standard for interconnection and interoperability of distributed energy resources with associated electric power systems interfaces. In: IEEE Std 1547-2018 (Revision of IEEE Std 1547-

[39] Xiang D, Turu JC, Muratel SM, Wang T. On-site LVRT testing method for full-power converter wind turbines. IEEE Transactions on Sustainable

2018;**54**(2):1663-1670

2003). 2018. pp. 1-138

Energy. 2017;**8**(1):395-403

*Development and Control of Generator-Converter Topology for Direct-Drive Wind Turbines DOI: http://dx.doi.org/10.5772/intechopen.85877*

[35] Singh A, Benzaquen J, Mirafzal B. Current source generator–converter topology for direct-drive wind turbines. IEEE Transactions on Industry Applications. March-April 2018;**54**(2):1663-1670

*Rotating Machinery*

region. IEEE Transactions on Power Electronics. 2016;**30**(10):6980-6991

[20] Tenca P, Rockhill AA, Lipo TA, Tricoli P. Current source topology for wind turbines with decreased mains current harmonics, further reducible via functional minimization. IEEE Transactions on Power Electronics.

[21] Lee JS, Lee KB, Blaabjerg F. Openswitch fault detection method of a back-to-back converter using NPC topology for wind turbine systems. IEEE Transactions on Industry Applications.

[22] Blaabjerg F, Ma K. Future on power electronics for wind turbine systems. IEEE Journal of Emerging and Selected Topics in Power Electron.

[23] Zavvos A, Mcdonald A, Mueller M. Optimisation tools for large permanent magnet generators for direct drive wind turbines. IET Renewable Power Generation. 2013;**7**(2):163-171

[24] Zhang Z, Zhao Y, Qiao W, Qu L. A space-vector-modulated sensorless direct-torque control for directdrive PMSG wind turbines. IEEE Transactions on Industry Applications.

Transactions on Industry Applications.

[26] Singh A, Mirafzal B. Three-phase single-stage boost inverter for direct drive wind turbines. In: 2016 IEEE

2008;**23**(3):1143-1155

2015;**51**(1):325-335

2013;**1**(3):139-151

2014;**50**(4):2331-2441

2015;**51**(4):2949-2958

[25] Potgieter JHJ, Kamper MJ. Design optimization of directly grid-connected PM machines for wind energy applications. IEEE

[19] Dai J, Xu DD, Wu B. A novel control scheme for current-source converterbased PMSG wind energy conversion systems. IEEE Transactions on Power Electronics. 2009;**24**(4):963-972

Energy Conversion Congress and Exposition (ECCE). Milwaukee, WI;

[27] Singh A, Mirafzal B. A low-voltage generator-converter topology for direct drive wind turbines. In: 2016 IEEE 7th International Symposium on Power Electron. For Distributed Gener. Sys. (PEDG). Vancouver, BC; 2016. pp. 1-6

[28] Wang Z, Wu B, Xu D, Zargari NR. A current-source-converter-based high-power high-speed PMSM drive with 420-Hz switching frequency. IEEE Transactions on Industrial Electronics.

[29] Mirafzal B, Saghaleini M, Kaviani A. An SVPWM-based switching pattern for stand-alone and grid-connected three-phase single-stage boost inverters. IEEE Transactions on Power Electronics.

[30] Singh A, Milani AA, Mirafzal B.

[31] Singh A, Kaviani AK, Mirafzal B. On dynamics models and stability of three-phase phasor PWM based CSI for stand-alone applications. IEEE Transactions on Industrial Electronics.

[32] Holmes DG, Lipo TA. Pulse Width Modulation for Power Converters: Principles and Practice. First ed.

New York, USA: John Wiley & Sons; 2003

[33] Fitzgerald AE, Kingsley C Jr, Umans SD. Electric Machinery. Sixth ed. New York, USA: McGraw Hill; 2003

[34] Northern Power NW 1500 Direct-Drive Generator –Report, [Online] Available from: http://www.nrel.gov/

docs/fy08osti/40177.pdf

Modified phasor pulse width modulation method for three-phase single-stage boost inverter. In: 2014 IEEE Applied Power Electron. Conf. And Expo. - APEC 2014. Fort Worth,

TX; 2014. pp. 1276-1280

May 2015;**62**(5):2698-2707

2012;**59**(7):2970-2981

2011;**26**(4):1102-1111

2016. pp. 1-7

**60**

[36] Fateh F, White WN, Gruenbacher D. A maximum power tracking technique for grid-connected DFIGbased wind turbines. IEEE Journal of Emerging and Selected Topics in Power Electronics. 2015;**3**(4):957-966

[37] Yuan X, Li Y. Control of variable pitch and variable speed direct-drive wind turbines in weak grid systems with active power balance. IET Renewable Power Generation. 2014;**8**(2):119-131

[38] IEEE standard for interconnection and interoperability of distributed energy resources with associated electric power systems interfaces. In: IEEE Std 1547-2018 (Revision of IEEE Std 1547- 2003). 2018. pp. 1-138

[39] Xiang D, Turu JC, Muratel SM, Wang T. On-site LVRT testing method for full-power converter wind turbines. IEEE Transactions on Sustainable Energy. 2017;**8**(1):395-403

Section 3

Wear and Abrasion

of Turbo Machines

63

Section 3
