# **New Advances in Melanoma**

#### Edited by Lionel Larribère

Printed Edition of the Special Issue Published in *Journal of Clinical Medicine*

www.mdpi.com/journal/jcm

**New Advances in Melanoma**


**Acknowledgements**

We would like to thank TUBITAK ULAKBIM, High Performance and GridComputing Center (TR-Grid e-Infrastructure) for the calculations reported in the theoretical part of this paper.

*Density Functional Theory - Recent Advances, New Perspectives and Applications*


*4.2.1 Influence of slot-pole combinations on pulsating torque*

Apart from the torque density, pulsating torque is also very important because a large pulsating torque will increase the vibration and noise of machines. **Figure 13** shows the cogging torque and ripple torque waveforms of 13-, 14-, 16-, 17-, and 19-rotor-slot FRPMMs. The stator slot number of these models is all chosen as 12. For the rated torque, we can see in **Figure 13(b)** that the 14-rotor-slot FRPMM yields the largest among the five models. As for the pulsating torque, we can see that the cogging torque and ripple torque of 16-rotor-slot FRPMM are the largest, and that of 19-rotor-slot FRPMM is the least. This phenomenon is related to the least common multiple of stator slot number and rotor slot number. The larger least common multiple, the lower pulsating torque. The least common multiples of the 13-, 14-, 16-, 17-, and 19-rotor-slot FRPMMs are 156, 84, 48, 204, and 228, respectively. Therefore, the 19-rotor-slot FRPMM exhibit the lowest cogging torque and ripple torque. However, attentions should be paid to use odd rotor number because it will cause other problems such as eccentricity stress. **Figure 14** compares the radial stress of the five FRPMM models. It can be seen that for the even rotor slot number FRPMMs, that is, 14 and 16 rotor slots, the stress harmonics only have even orders, which will not lead to eccentricity. However, for the odd rotor slot number

FRPMMs, that is, 13, 17, and 19 rotor slots, there are many odd stress harmonics. Since the first-order harmonic is dominant for the eccentricity, the 13-rotor-slot

recommended. The first-order stress harmonic for 17 and 19 rotor slots are very

The influences of split ratio and PM thickness on cogging torque and ripple torque of FRPMMs are also analyzed in **Figure 15**. This figure is plotted based on the 14-rotor-slot, which is chosen because it has the largest torque density and a relatively low pulsating torque, as shown in **Figure 13**. It can be found in **Figure 15(a)** that the cogging torque increases with the PM thickness and the split ratio. When the PM thickness increases, the airgap flux density increases, and thus the interaction between the PMs and slot-teeth becomes greater, which leads to a higher cogging torque. As the split ratio increases, the airgap radius increases, hence the cogging torque increases with the split ratio [30]. As for the ripple torque, the ripple torque has the maximum value when the split ratio is around 0.66. This is because the ripple torque is not only related to the slot structure but also influenced by the electric loading. As aforementioned, the pulsating torque resulting from the slot structure is increased with the split ratio. However, as the split ratio increases, the slot area is reduced and the electric loading gets smaller and smaller, so the ripple torque resulting from the electric loading becomes lower. Considering these two impacts, the ripple torque has a maximal value when the split ratio changes.

As we know, the airgap structure is significant for the pulsating torque because

the pulsating torque results from the interaction between the two sides of the airgap, that is, stator and rotor. Therefore, this chapter also analyzes the influences of stator slot opening ratio and rotor slot opening ratio on cogging torque and ripple torque. Here, the stator/rotor slot opening ratio is defined as the ratio of stator/rotor slot opening width to the stator/rotor slot pitch. **Figure 16** shows the variation of cogging torque and ripple torque with the two slot opening ratios. It can be seen that the cogging torque increases with the stator slot opening ratio. The reason is that a larger stator slot opening ratio reduces the PM width and the smoothness of PM MMF, thus the changing of the PM MMF along the tangential direction increases

*Effect of split ratio and PM thickness on pulsating torque performances: (a) cogging torque (%); (b) ripple*

FRPMM has a large eccentricity stress. Therefore, 13-rotor-slot is not

*4.2.2 Influence of PM thickness and split ratio on pulsating torque*

*4.2.3 Influence of slot opening ratios on pulsating torque*

**Figure 15.**

*torque (%).*

**83**

small, so their eccentricity can be neglected.

*DOI: http://dx.doi.org/10.5772/intechopen.92428*

*Flux Reversal Machine Design*

**Figure 13.** *Effect of slot-pole combination on pulsating torque performances: (a) cogging torque waveforms (%); (b) rated torque waveforms.*

**Figure 14.**

*Radial stress analysis of the FRPMMs: (a) 13-rotor-slot; (b) 14-rotor-slot; (c) 16-rotor-slot; (d) 17-rotor-slot; (e) 19-rotor-slot.*

#### *Flux Reversal Machine Design DOI: http://dx.doi.org/10.5772/intechopen.92428*

FRPMMs, that is, 13, 17, and 19 rotor slots, there are many odd stress harmonics. Since the first-order harmonic is dominant for the eccentricity, the 13-rotor-slot FRPMM has a large eccentricity stress. Therefore, 13-rotor-slot is not recommended. The first-order stress harmonic for 17 and 19 rotor slots are very small, so their eccentricity can be neglected.

#### *4.2.2 Influence of PM thickness and split ratio on pulsating torque*

The influences of split ratio and PM thickness on cogging torque and ripple torque of FRPMMs are also analyzed in **Figure 15**. This figure is plotted based on the 14-rotor-slot, which is chosen because it has the largest torque density and a relatively low pulsating torque, as shown in **Figure 13**. It can be found in **Figure 15(a)** that the cogging torque increases with the PM thickness and the split ratio. When the PM thickness increases, the airgap flux density increases, and thus the interaction between the PMs and slot-teeth becomes greater, which leads to a higher cogging torque. As the split ratio increases, the airgap radius increases, hence the cogging torque increases with the split ratio [30]. As for the ripple torque, the ripple torque has the maximum value when the split ratio is around 0.66. This is because the ripple torque is not only related to the slot structure but also influenced by the electric loading. As aforementioned, the pulsating torque resulting from the slot structure is increased with the split ratio. However, as the split ratio increases, the slot area is reduced and the electric loading gets smaller and smaller, so the ripple torque resulting from the electric loading becomes lower. Considering these two impacts, the ripple torque has a maximal value when the split ratio changes.

#### *4.2.3 Influence of slot opening ratios on pulsating torque*

As we know, the airgap structure is significant for the pulsating torque because the pulsating torque results from the interaction between the two sides of the airgap, that is, stator and rotor. Therefore, this chapter also analyzes the influences of stator slot opening ratio and rotor slot opening ratio on cogging torque and ripple torque. Here, the stator/rotor slot opening ratio is defined as the ratio of stator/rotor slot opening width to the stator/rotor slot pitch. **Figure 16** shows the variation of cogging torque and ripple torque with the two slot opening ratios. It can be seen that the cogging torque increases with the stator slot opening ratio. The reason is that a larger stator slot opening ratio reduces the PM width and the smoothness of PM MMF, thus the changing of the PM MMF along the tangential direction increases

*Effect of split ratio and PM thickness on pulsating torque performances: (a) cogging torque (%); (b) ripple torque (%).*

**Figure 13.**

**Figure 14.**

**82**

*(e) 19-rotor-slot.*

*torque waveforms.*

*Direct Torque Control Strategies of Electrical Machines*

*Effect of slot-pole combination on pulsating torque performances: (a) cogging torque waveforms (%); (b) rated*

*Radial stress analysis of the FRPMMs: (a) 13-rotor-slot; (b) 14-rotor-slot; (c) 16-rotor-slot; (d) 17-rotor-slot;*

**Figure 16.**

*Effect of stator slot opening ratio and rotor slot opening ratio on pulsating torque: (a) cogging torque (%); (b) ripple torque (%).*

the cogging torque. As for the rotor slot opening ratio, which simultaneously influences all the harmonic contents of the airgap permeance, it has great and nonlinear impact on the pulsating torque. Since the pulsating torque results from the interaction of multi permeance harmonics, the variation of pulsating torque changes nonlinearly with the rotor slot opening ratio. It can be seen in **Figure 16** that the optimal cogging torque and ripple torque can be achieved when the stator slot opening ratio and rotor slot opening ratio are around 0.25 and 0.7, respectively.

#### **4.3 Power factor performances**

#### *4.3.1 Influence of stator inner diameter and PM thickness on power factor*

Since the power factor of FRPMMs is usually low, which is around 0.4–0.7, meanwhile a low power factor will increase the converter capacity and cost, the influences of key parameters on the power factor should be also analyzed to achieve a relatively high power factor. The power factor can be given as:

$$PF = \mathbf{1} / \sqrt{\mathbf{1} + \left(\frac{L\_t I\_s}{\mu\_m}\right)^2} \tag{31}$$

*4.3.2 Influence of slot opening ratio on power factor*

*Effect of PM thickness/airgap length on power factor.*

*Effect of stator inner diameter/airgap length on power factor.*

**Figure 17.**

*Flux Reversal Machine Design*

*DOI: http://dx.doi.org/10.5772/intechopen.92428*

**Figure 18.**

**85**

optimal value for the stator slot opening ratio.

Another important parameter that influences the airgap structure is the slot opening ratio. Hence, **Figures 19** and **20** analyzes the effect of stator slot opening ratio and rotor slot opening ratio on pulsating torque performances, respectively. It can be seen in **Figure 19** that the maximum power factor can be obtained when the stator slot opening ratio is approximately to 0.3. The explanation is as follows. When the stator slot opening ratio is too small, the slot leakage flux between the stator tips is large, thus the main flux is reduced, and the back-EMF is lowered, resulting in smaller back-EMF. And when the stator slot opening ratio is too large, the PM width will be narrower. Although the slot leakage flux is reduced, the main flux is not high due to the narrower PMs, thus the back-EMF is lowered. Therefore, the stator slot opening ratio cannot be too small or too large, that is, there is an

Then, the influences of rotor slot opening ratio on power factor can be seen in **Figure 20**. It indicates that when the rotor slot opening ratio is around 0.7, the

where *Is* is the winding current, *Ls* is the synchronous inductance (because the saliency ratio is approximate to 1, *Ld* ≈ *Lq*), and *ψm* is the PM flux linkage. Then, the effect of stator inner diameter on power factor is shown in **Figure 17**. Here, the stator outer diameter is kept as 124 mm, and the airgap length is fixed as 0.5 mm. It can be found that with the increase of stator inner diameter, the power factor increases continuously. The reason is that with the increase of stator inner diameter, the slot area decreases, so the winding turns per phase decreases, thus leading to the reduction of the synchronous inductance *Ls*. The lower *Ls*, the higher power factor, as shown in Eq. (31). Apart from the stator inner diameter, another important parameter affecting the power factor is the PM thickness *hm*. **Figure 18** investigates the variation of power factor with respect to the PM thickness. It indicates that the power factor initially increases with the PM thickness but then decreases. The reason is explained as follows. As the PM thickness increases, the PM flux linkage *ψm* becomes larger, so the power factor increases. However, the synchronous inductance *Ls* also increases with the PM thickness, which leads to the reduction of power factor afterwards. Therefore, there is an optimal PM thickness for a maximum achievable power factor.

*Flux Reversal Machine Design DOI: http://dx.doi.org/10.5772/intechopen.92428*

**Figure 17.** *Effect of stator inner diameter/airgap length on power factor.*

**Figure 18.** *Effect of PM thickness/airgap length on power factor.*

#### *4.3.2 Influence of slot opening ratio on power factor*

Another important parameter that influences the airgap structure is the slot opening ratio. Hence, **Figures 19** and **20** analyzes the effect of stator slot opening ratio and rotor slot opening ratio on pulsating torque performances, respectively. It can be seen in **Figure 19** that the maximum power factor can be obtained when the stator slot opening ratio is approximately to 0.3. The explanation is as follows. When the stator slot opening ratio is too small, the slot leakage flux between the stator tips is large, thus the main flux is reduced, and the back-EMF is lowered, resulting in smaller back-EMF. And when the stator slot opening ratio is too large, the PM width will be narrower. Although the slot leakage flux is reduced, the main flux is not high due to the narrower PMs, thus the back-EMF is lowered. Therefore, the stator slot opening ratio cannot be too small or too large, that is, there is an optimal value for the stator slot opening ratio.

Then, the influences of rotor slot opening ratio on power factor can be seen in **Figure 20**. It indicates that when the rotor slot opening ratio is around 0.7, the

the cogging torque. As for the rotor slot opening ratio, which simultaneously influences all the harmonic contents of the airgap permeance, it has great and nonlinear impact on the pulsating torque. Since the pulsating torque results from the interaction of multi permeance harmonics, the variation of pulsating torque changes nonlinearly with the rotor slot opening ratio. It can be seen in **Figure 16** that the optimal cogging torque and ripple torque can be achieved when the stator slot opening ratio and rotor slot opening ratio are around 0.25 and 0.7, respectively.

*Effect of stator slot opening ratio and rotor slot opening ratio on pulsating torque: (a) cogging torque (%);*

Since the power factor of FRPMMs is usually low, which is around 0.4–0.7, meanwhile a low power factor will increase the converter capacity and cost, the influences of key parameters on the power factor should be also analyzed to achieve

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi

*LsIs ψ m* � �2

(31)

1 þ

where *Is* is the winding current, *Ls* is the synchronous inductance (because the saliency ratio is approximate to 1, *Ld* ≈ *Lq*), and *ψm* is the PM flux linkage. Then, the effect of stator inner diameter on power factor is shown in **Figure 17**. Here, the stator outer diameter is kept as 124 mm, and the airgap length is fixed as 0.5 mm. It can be found that with the increase of stator inner diameter, the power factor increases continuously. The reason is that with the increase of stator inner diameter, the slot area decreases, so the winding turns per phase decreases, thus leading to the reduction of the synchronous inductance *Ls*. The lower *Ls*, the higher power factor, as shown in Eq. (31). Apart from the stator inner diameter, another important parameter affecting the power factor is the PM thickness *hm*. **Figure 18** investigates the variation of power factor with respect to the PM thickness. It indicates that the power factor initially increases with the PM thickness but then decreases. The reason is explained as follows. As the PM thickness increases, the PM flux linkage *ψm* becomes larger, so the power factor increases. However, the synchronous inductance *Ls* also increases with the PM thickness, which leads to the reduction of power factor afterwards. Therefore, there is an optimal PM thickness for a maxi-

s

*4.3.1 Influence of stator inner diameter and PM thickness on power factor*

a relatively high power factor. The power factor can be given as:

*PF* ¼ 1*=*

**4.3 Power factor performances**

*Direct Torque Control Strategies of Electrical Machines*

**Figure 16.**

*(b) ripple torque (%).*

mum achievable power factor.

**84**

chapter. Since the magnetic properties of PM materials are sensitive to temperature,

**Figure 21** shows the demagnetization curve of the magnets. The upper half is a straight line, and lower half under the knee point *Bknee* is a curved line. When the FRPMM works on the straight line (such as point *P*1), the return line coincides with the demagnetization curve, and the magnetic performance of the magnets will not be lost. However, when the armature equivalent MMF *Ha*´ is too large at load condition, or the knee point is too high, the working point *Bknee* is moved to *P*2. At this time, the recovery line does not coincide with the original demagnetization line, thus the intersection of the *B*-axis changes from *Br* to *Br*1, causing the irreversible demagnetization. Then, the PM properties and machine performances will no longer return to the original. So, the PM flux density should be examined in order to check the risk of irreversible demagnetization. As we know, the PM flux density is determined by the design parameters such as electric loading *Ae*, PM thickness *hm*, rotor slot opening ratio, etc. So, in this chapter, the effects of electric loading *Ae*, PM thickness *hm*, rotor slot opening ratio *bo/t* on PM demagnetization performances of FRPMMs will be studied. For instance, the PM material is selected as N38SH, and

**Figure 22** shows the PM flux density of a 12-stator-slot/14-rotor-slot FRPMM when the electric loading *Ae* is 1600A/cm, the PM thickness *hm* is 3 mm, rotor slot opening ratio *bo/t* is 0.65. It can be seen that the PM flux density distribution varies with the rotor position. When the rotor position is 140°, the PM does not demagnetize, while at 0° and 340°, the PM will demagnetize. Hence, in the following analysis, the PM flux density at the most severe moment of demagnetization is

**Figure 23** studies the magnetic flux density distribution in the PMs under dif-

ferent electric loadings. It can be found that the larger electric loading *Ae*, the smaller minimum flux density. This is because the larger electric loading, the higher armature MMF *Ha*´, and the more left operating point *P*2, so the lower flux density

. When

and the temperature coefficient of NdFeB magnet is as high as 0.126%K1

tization performances of FRPMMs at different conditions.

*Flux Reversal Machine Design*

*DOI: http://dx.doi.org/10.5772/intechopen.92428*

knee point of the PM flux density at 100°C is 0.35 T.

selected.

**Figure 21.**

**87**

*PM demagnetization curve.*

the current of FRPMMs is large, the winding heating can easily affect the PMs attached to the stator teeth surface, causing the decrease of PM magnetic performances. On the other hand, when the winding current is large, the demagnetizing effect of the armature field is enhanced, and thus the PMs have the possibility to be demagnetized. Therefore, it is of great importance to investigate the PM demagne-

**Figure 19.** *Effect of stator slot opening ratio on power factor.*

**Figure 20.** *Effect of rotor slot opening ratio on power factor.*

power factor reaches the maximal value. This is because the power factor is mainly influenced by the back-EMF. When the rotor slot opening ratio increases, the effective airgap length becomes smaller, thus the main flux is increased and the back-EMF is improved. As a result, the power factor is increased. When the rotor slot opening ratio keeps increasing, the flux modulation effect of the rotor teeth becomes weaker and weaker, thus the smaller modulated flux, and the lower back-EMF. Therefore, there is also an optimal value for rotor slot opening ratio when a high power factor is demanded.

#### **4.4 PM demagnetization performances**

For PM machines, PM demagnetization performances are very important because it is highly related to the safe operation and machine reliability. Therefore, the PM demagnetization performances of FRPMMs should be analyzed in this

#### *Flux Reversal Machine Design DOI: http://dx.doi.org/10.5772/intechopen.92428*

chapter. Since the magnetic properties of PM materials are sensitive to temperature, and the temperature coefficient of NdFeB magnet is as high as 0.126%K1 . When the current of FRPMMs is large, the winding heating can easily affect the PMs attached to the stator teeth surface, causing the decrease of PM magnetic performances. On the other hand, when the winding current is large, the demagnetizing effect of the armature field is enhanced, and thus the PMs have the possibility to be demagnetized. Therefore, it is of great importance to investigate the PM demagnetization performances of FRPMMs at different conditions.

**Figure 21** shows the demagnetization curve of the magnets. The upper half is a straight line, and lower half under the knee point *Bknee* is a curved line. When the FRPMM works on the straight line (such as point *P*1), the return line coincides with the demagnetization curve, and the magnetic performance of the magnets will not be lost. However, when the armature equivalent MMF *Ha*´ is too large at load condition, or the knee point is too high, the working point *Bknee* is moved to *P*2. At this time, the recovery line does not coincide with the original demagnetization line, thus the intersection of the *B*-axis changes from *Br* to *Br*1, causing the irreversible demagnetization. Then, the PM properties and machine performances will no longer return to the original. So, the PM flux density should be examined in order to check the risk of irreversible demagnetization. As we know, the PM flux density is determined by the design parameters such as electric loading *Ae*, PM thickness *hm*, rotor slot opening ratio, etc. So, in this chapter, the effects of electric loading *Ae*, PM thickness *hm*, rotor slot opening ratio *bo/t* on PM demagnetization performances of FRPMMs will be studied. For instance, the PM material is selected as N38SH, and knee point of the PM flux density at 100°C is 0.35 T.

**Figure 22** shows the PM flux density of a 12-stator-slot/14-rotor-slot FRPMM when the electric loading *Ae* is 1600A/cm, the PM thickness *hm* is 3 mm, rotor slot opening ratio *bo/t* is 0.65. It can be seen that the PM flux density distribution varies with the rotor position. When the rotor position is 140°, the PM does not demagnetize, while at 0° and 340°, the PM will demagnetize. Hence, in the following analysis, the PM flux density at the most severe moment of demagnetization is selected.

**Figure 23** studies the magnetic flux density distribution in the PMs under different electric loadings. It can be found that the larger electric loading *Ae*, the smaller minimum flux density. This is because the larger electric loading, the higher armature MMF *Ha*´, and the more left operating point *P*2, so the lower flux density

**Figure 21.** *PM demagnetization curve.*

power factor reaches the maximal value. This is because the power factor is mainly influenced by the back-EMF. When the rotor slot opening ratio increases, the effective airgap length becomes smaller, thus the main flux is increased and the back-EMF is improved. As a result, the power factor is increased. When the rotor slot opening ratio keeps increasing, the flux modulation effect of the rotor teeth becomes weaker and weaker, thus the smaller modulated flux, and the lower back-EMF. Therefore, there is also an optimal value for rotor slot opening ratio when a

For PM machines, PM demagnetization performances are very important because it is highly related to the safe operation and machine reliability. Therefore, the PM demagnetization performances of FRPMMs should be analyzed in this

high power factor is demanded.

*Effect of rotor slot opening ratio on power factor.*

**Figure 19.**

**Figure 20.**

**86**

*Effect of stator slot opening ratio on power factor.*

*Direct Torque Control Strategies of Electrical Machines*

**4.4 PM demagnetization performances**

**Figure 22.**

*PM demagnetization at different rotor positions: (a) rotor position = 0°; (b) rotor position = 140°; (c) rotor position = 340°.*

**Figure 23.** *Influence of* Ae *on PM demagnetization.*

in the magnets. When the electric loading *Ae* is 1400A/cm, the PM irreversible demagnetization just occurs. In addition, it can be seen that the entire magnetic flux density map is skewed to the right. This is because the N-pole magnet is intercepted in this analysis, and there is an S-pole magnet next to the N-pole magnet. There is PM pole leakage flux between the S-pole magnet (negative axis) and the N-pole magnet (positive axis), so the magnetic flux density around the 0 position is lower, and away from the 0 position, the magnetic flux density gradually rises.

and the PM thickness is selected as five times the airgap length, that is, 3 mm. The larger rotor slot opening ratio, the narrower rotor teeth, thus the more saturated rotor teeth, and the smaller magnetic reluctance. As shown in **Figure 20**, when the magnetic gets smaller, the more left operating point *P*2, and thus the lower PM flux density. It can be seen in **Figure 25** that when the rotor slot opening ratio *bo*/*t* is 0.9, the irreversible PM demagnetization just occurs. In Ref. [28], it is claimed that the maximum back-EMF can be achieved when the rotor slot opening ratio *bo*/*t* is around 0.6. So, during the design process, the optimal rotor slot opening ratio can

be directly applied without consideration of the PM demagnetization risk.

The geometrical parameters of stator and rotor are shown in **Figure 26**. The

**5. Geometric design of stator and rotor**

no-load flux of each winding pole could be calculated as:

**5.1 Stator design**

**89**

**Figure 24.**

**Figure 25.**

*Influence of* hm *on PM demagnetization.*

*Flux Reversal Machine Design*

*DOI: http://dx.doi.org/10.5772/intechopen.92428*

*Influence of* bo*/*t *on PM demagnetization.*

**Figure 24** analyzes the effect of PM thickness *hm* on the PM demagnetization performances. At this time, the electric loading is chosen as 800 A/cm, and the rotor slot opening ratio is selected as 0.65. It can be seen in **Figure 24** that when the PM thickness *hm* is less than 2.5 mm, the irreversible demagnetization will happen, while when the PM thickness *hm* is larger than 2.5 mm, the irreversible demagnetization will not. In this model, the airgap length is 0.5 mm. Therefore, in the design stage, the PM thickness should be better to set as five times or more the airgap length. Considering the back-EMF, it is claimed in [3] that when the PM thickness is about three times the airgap length, the back-EMF will reach the maximum. But considering both back-EMF and PM demagnetization risk, it is safer to set the PM thickness as about five times airgap length.

**Figure 25** shows the influences of rotor slot opening ratio *bo*/*t* on the flux density distribution inside the PMs. At this time, the electric loading is chosen as 800 A/cm,

**Figure 24.** *Influence of* hm *on PM demagnetization.*

**Figure 25.** *Influence of* bo*/*t *on PM demagnetization.*

and the PM thickness is selected as five times the airgap length, that is, 3 mm. The larger rotor slot opening ratio, the narrower rotor teeth, thus the more saturated rotor teeth, and the smaller magnetic reluctance. As shown in **Figure 20**, when the magnetic gets smaller, the more left operating point *P*2, and thus the lower PM flux density. It can be seen in **Figure 25** that when the rotor slot opening ratio *bo*/*t* is 0.9, the irreversible PM demagnetization just occurs. In Ref. [28], it is claimed that the maximum back-EMF can be achieved when the rotor slot opening ratio *bo*/*t* is around 0.6. So, during the design process, the optimal rotor slot opening ratio can be directly applied without consideration of the PM demagnetization risk.

### **5. Geometric design of stator and rotor**

#### **5.1 Stator design**

The geometrical parameters of stator and rotor are shown in **Figure 26**. The no-load flux of each winding pole could be calculated as:

in the magnets. When the electric loading *Ae* is 1400A/cm, the PM irreversible demagnetization just occurs. In addition, it can be seen that the entire magnetic flux density map is skewed to the right. This is because the N-pole magnet is intercepted in this analysis, and there is an S-pole magnet next to the N-pole magnet. There is PM pole leakage flux between the S-pole magnet (negative axis) and the N-pole magnet (positive axis), so the magnetic flux density around the 0 position is lower,

*PM demagnetization at different rotor positions: (a) rotor position = 0°; (b) rotor position = 140°; (c) rotor*

*Direct Torque Control Strategies of Electrical Machines*

**Figure 24** analyzes the effect of PM thickness *hm* on the PM demagnetization performances. At this time, the electric loading is chosen as 800 A/cm, and the rotor slot opening ratio is selected as 0.65. It can be seen in **Figure 24** that when the PM thickness *hm* is less than 2.5 mm, the irreversible demagnetization will happen, while when the PM thickness *hm* is larger than 2.5 mm, the irreversible demagnetization will not. In this model, the airgap length is 0.5 mm. Therefore, in the design stage, the PM thickness should be better to set as five times or more the airgap length. Considering the back-EMF, it is claimed in [3] that when the PM thickness is about three times the airgap length, the back-EMF will reach the maximum. But considering both back-EMF and PM demagnetization risk, it is safer to set the PM

**Figure 25** shows the influences of rotor slot opening ratio *bo*/*t* on the flux density distribution inside the PMs. At this time, the electric loading is chosen as 800 A/cm,

and away from the 0 position, the magnetic flux density gradually rises.

thickness as about five times airgap length.

**Figure 22.**

**Figure 23.**

**88**

*Influence of* Ae *on PM demagnetization.*

*position = 340°.*

*Direct Torque Control Strategies of Electrical Machines*

$$
\phi\_m = \mathcal{Z}\lambda\_w l\_{\text{stk}} B\_m / \pi \tag{32}
$$

Then, next step is to calculate the stator outer radius *ro*. Firstly, the total slot area of all the stator slots *Aslot* can be written based on the winding electric loading *Ae*

where *Sfg* is the slot fill factor. Meanwhile, the total slot area of all the stator slots

Combining the Eqs. (39) and (40), the slot depth *hs* can be determined. Then,

Defining the average flux density of each rotor yoke and middle of rotor tooth as *Bry*, and *Brt*, respectively, the rotor yoke thickness *hry* and rotor tooth width *wrt* are able to be achieved using the similar derivation procedure as Eq. (35) and Eq. (36).

Based on the analytical equations and the investigations of key performances in the former parts, a quick and accurate analytical design of a FRPMM can be realized

1.Based on the performance investigations in **Figures 9–24**, the initial design values, including combination of stator slot and rotor slot number, electric loading, equivalent magnetic loading, airgap length, materials of active parts,

2.Then, assuming an appropriate aspect ratio *klr*, the airgap radius *rg,* and the

3.Based on Eqs. (32–44), the detailed geometric parameters of the stator core and the rotor core are able to be obtained. Therefore, the stator outer diameter

4.After that, check if the stator outer diameter and the machine total length satisfy the required design specifications. If so, proceed to the FEA

stack length *lstk* can be worked out using Eqs. (29) and (30).

*ro* and machine total length *lo* can be finally determined.

*Ae=JeSfg* (39)

2 � *Zswths* (40)

*ro* ¼ *rg* þ *hm* þ *h*1 þ *hs* þ *hy* (41)

*hry* ¼ *rgBm=ZrBrykstk* (42) *wrt* ¼ 4*rgBm=ZrkstkBrt* (43)

*hrs* ¼ *rs* þ *hry* (44)

*Aslot* ¼ 2*π rg* þ *hm*

2 � *π rg* þ *hm* þ *h*1

*Aslot* can be also derived out using the structural parameters:

*Aslot* ¼ *π rg* þ *hm* þ *h*1 þ *hs*

the stator outer radius *ro* can be given as:

Finally, the *hry* and *wrt* are given as:

Then, the rotor slot depth *hrs* is determined as:

by following these procedures (as depicted in **Figure 27**):

**6. Design methodology and evaluations**

etc. can be firstly selected.

and the current density *Je*:

*Flux Reversal Machine Design*

*DOI: http://dx.doi.org/10.5772/intechopen.92428*

**5.2 Rotor design**

**6.1 Design procedure**

**91**

where *λw* is the winding pitch. If full-pitch winding is adopted, the winding pitch is able to be written as:

$$
\lambda\_w = 2\pi r\_\text{g} / 2P \tag{33}
$$

Then, the no-load flux of each winding pole *ϕm* in Eq. (32) could change to:

$$
\phi\_m = 2r\_\text{g} l\_{stk} B\_m / P \tag{34}
$$

Defining the average flux density at the stator yoke as *By*, the stator yoke thickness *hy* can therefore be deduced as:

$$h\_{\gamma} = \frac{\phi\_m}{2B\_{\gamma}k\_{stk}l\_{stk}} = \frac{r\_{\text{g}}B\_m}{PB\_{\gamma}k\_{stk}}\tag{35}$$

Similarly, defining the average flux density at the middle of stator tooth as *Bt*, the stator tooth width is able to be worked out:

$$w\_t = \frac{\phi\_m}{\text{3SPP}k\_{stk}l\_{stk}B\_t} = \frac{4r\_{\text{g}}B\_m}{Z\_s k\_{stk}B\_t} \tag{36}$$

Moreover, in order to simultaneously maintain a relatively large torque density as well as reduce the risk of PM demagnetization, the PM thickness is recommended to be:

$$h\_m = 4\text{g} \sim 6\text{g} \tag{37}$$

where *g* is the airgap length. Since the optimal torque density is often obtained when the slot opening ratio is approximate to 0.25 [28], the stator slot opening width *wo* could be written as:

$$
\omega\_o = \pi (r\_\text{g} + h\_m) / 2Z\_s \tag{38}
$$

**Figure 26.** *Geometry of stator and rotor.*

*Flux Reversal Machine Design DOI: http://dx.doi.org/10.5772/intechopen.92428*

Then, next step is to calculate the stator outer radius *ro*. Firstly, the total slot area of all the stator slots *Aslot* can be written based on the winding electric loading *Ae* and the current density *Je*:

$$A\_{\rm slot} = 2\pi (r\_{\rm g} + h\_m) A\_{\rm e} / J\_{\rm e} \text{S}\_{\rm f\bar{\rm g}} \tag{39}$$

where *Sfg* is the slot fill factor. Meanwhile, the total slot area of all the stator slots *Aslot* can be also derived out using the structural parameters:

$$A\_{slat} = \pi \left(r\_{\rm g} + h\_m + h\_1 + h\_t\right)^2 - \pi \left(r\_{\rm g} + h\_m + h\_1\right)^2 - Z\_s w\_l h\_s \tag{40}$$

Combining the Eqs. (39) and (40), the slot depth *hs* can be determined. Then, the stator outer radius *ro* can be given as:

$$r\_o = r\_\text{g} + h\_m + h\_1 + h\_s + h\_y \tag{41}$$

#### **5.2 Rotor design**

*ϕm* ¼ 2*λwlstkBm=π* (32)

*λw* ¼ 2*πrg=*2*P* (33)

*ϕm* ¼ 2*rg lstkBm=P* (34)

*hm* ¼ 4*g* � 6*g* (37)

*=*2*Zs* (38)

(35)

(36)

where *λw* is the winding pitch. If full-pitch winding is adopted, the winding pitch

Then, the no-load flux of each winding pole *ϕm* in Eq. (32) could change to:

Defining the average flux density at the stator yoke as *By*, the stator yoke

Similarly, defining the average flux density at the middle of stator tooth as *Bt*,

Moreover, in order to simultaneously maintain a relatively large torque density as well as reduce the risk of PM demagnetization, the PM thickness is recommended

where *g* is the airgap length. Since the optimal torque density is often obtained when the slot opening ratio is approximate to 0.25 [28], the stator slot opening

*wo* ¼ *π rg* þ *hm*

3*SPPkstklstkBt*

¼ *rgBm PBykstk*

> ¼ 4*rgBm ZskstkBt*

*hy* ¼ *ϕm* 2*Bykstklstk*

*wt* ¼ *ϕm*

is able to be written as:

to be:

**Figure 26.**

**90**

*Geometry of stator and rotor.*

width *wo* could be written as:

thickness *hy* can therefore be deduced as:

*Direct Torque Control Strategies of Electrical Machines*

the stator tooth width is able to be worked out:

Defining the average flux density of each rotor yoke and middle of rotor tooth as *Bry*, and *Brt*, respectively, the rotor yoke thickness *hry* and rotor tooth width *wrt* are able to be achieved using the similar derivation procedure as Eq. (35) and Eq. (36). Finally, the *hry* and *wrt* are given as:

$$h\_{\gamma\gamma} = r\_{\rm g} B\_m / Z\_r B\_{r\rm ry} k\_{\rm stk} \tag{42}$$

$$
\omega\_{rt} = 4r\_{\rm g}B\_m / Z\_r k\_{stk} B\_{rt} \tag{43}
$$

Then, the rotor slot depth *hrs* is determined as:

$$h\_{\pi\circ} = r\_{\circ} + h\_{r\circ} \tag{44}$$

### **6. Design methodology and evaluations**

#### **6.1 Design procedure**

Based on the analytical equations and the investigations of key performances in the former parts, a quick and accurate analytical design of a FRPMM can be realized by following these procedures (as depicted in **Figure 27**):

verifications of machine performances. If not, reset the initial values such as combination of stator slot and rotor slot number, electric loading, equivalent magnetic loading, etc.

5.Conducting FEA simulations, the electromagnetic performances such as back-EMF, average torque, pulsating torque, power factor, efficiency, etc. can be obtained. Check if all the performances satisfy the design specifications. If not, adjust the design parameters in the former steps and iterate the design flow

In order show the effectiveness of the introduced analytical method, a FRPMM is designed based on the method. **Table 4** shows the specifications of the FRPMM, which mainly includes the rated torque, machine volume, cooling method, rated power, and speed. According to the rated torque, a design margin of 5% is suggested so as to make sure the torque output. Therefore, the requirement of the torque is 8.4 Nm for this design. Then, the combination of stator slots and rotor slots is determined in the first place. This combination is selected due to its high torque density and low pulsating torque, as shown in **Figure 13**. Then, since the cooling method is natural cooling, the electric loading and the equivalent magnetic loading are chosen as 300A/cm and 0.2 T, respectively. After that, based on the output torque value 8.4 Nm and Eq. (29), the airgap radius is determined as 38.5 mm. Furthermore, assuming the yoke flux density of stator core and rotor core as 1.0 T,

and the teeth flux density of stator core and rotor core as 1.2 T, the detailed

**Parameter Value Parameter Value** Rated torque 8 Nm Rated speed 300 rpm Rotor inner diameter 32 mm PM material N38SH Stator outer diameter 130 mm Stack length 120 mm Airgap length 0.6 mm Iron material 50WW470 Cooling method Natural cooling Rated power 0.25 kW

Stator Outer diameter 124 mm Inner diameter 79 mm

Magnet PM thickness 3 mm Magnet width 7.8 mm Rotor Outer diameter 77.8 mm Slot depth 10.4 mm

**Parameter Value Parameter Value**

Turns per phase 300 Teeth width 11.5 mm Slot number 12 Yoke thickness 6 mm Slot depth 13.5 mm Yoke flux density 1.0 T Winding pole pair 1 Teeth flux density 1.1 T

Teeth flux density 1.2 T Yoke thickness 12.5 mm Inner diameter 32 mm Yoke flux density 1.0 T Teeth width 4 mm Slot number 17

until every output meets the requirement.

6.Finally, it is the result output.

*DOI: http://dx.doi.org/10.5772/intechopen.92428*

*Flux Reversal Machine Design*

**6.2 Case study**

**Table 4.**

**Table 5.**

**93**

*Design specifications of a three-phase FRPMM.*

*Design parameters of the FRPMM using the design method.*

**Figure 27.** *Design flow of FRPMM.*

### *Flux Reversal Machine Design DOI: http://dx.doi.org/10.5772/intechopen.92428*


**Bulked Segregant Analysis Coupled with Whole-Genome Sequencing (BSA-Seq) Mapping Identifies a Novel** *pi21* **Haplotype Conferring Basal Resistance to Rice Blast Disease**

**Tingmin Liang 1,2,**†**, Wenchao Chi 1,**†**, Likun Huang 3, Mengyu Qu 1, Shubiao Zhang 3, Zi-Qiang Chen 2, Zai-Jie Chen 2, Dagang Tian 2, Yijie Gui 2, Xiaofeng Chen 1, Zonghua Wang 1,4, Weiqi Tang 1,\* and Songbiao Chen 1,\***

Received: 6 December 2019; Accepted: 19 March 2020; Published: 21 March 2020

**Abstract:** Basal or partial resistance has been considered race-non-specific and broad-spectrum. Therefore, the identification of genes or quantitative trait loci (QTLs) conferring basal resistance and germplasm containing them is of significance in breeding crops with durable resistance. In this study, we performed a bulked segregant analysis coupled with whole-genome sequencing (BSA-seq) to identify QTLs controlling basal resistance to blast disease in an F2 population derived from two rice varieties, 02428 and LiXinGeng (LXG), which differ significantly in basal resistance to rice blast. Four candidate QTLs, *qBBR-4*, *qBBR-7*, *qBBR-8*, and *qBBR-11*, were mapped on chromosomes 4, 7, 8, and 11, respectively. Allelic and genotypic association analyses identified a novel haplotype of the durable blast resistance gene *pi21* carrying double deletions of 30 bp and 33 bp in 02428 (*pi21-2428*) as a candidate gene of *qBBR-4*. We further assessed haplotypes of *Pi21* in 325 rice accessions, and identified 11 haplotypes among the accessions, of which eight were novel types. While the resistant *pi21* gene was found only in *japonica* before, three Chinese *indica* varieties, ShuHui881, Yong4, and ZhengDa4Hao, were detected carrying the resistant *pi21-2428* allele. The *pi21-2428* allele and *pi21-2428*-containing rice germplasm, thus, provide valuable resources for breeding rice varieties, especially *indica* rice varieties, with durable resistance to blast disease. Our results also lay the foundation for further identification and functional characterization of the other three QTLs to better understand the molecular mechanisms underlying rice basal resistance to blast disease.

**Keywords:** rice; blast disease; partial resistance; *pi21*; haplotype


**1. Introduction**

Hydroxyapatite (HAp, Ca10(PO4 )6 (OH)2 ) is one of the most usual forms of calcium phosphate and has the similar chemical composition to the mineral phase of bone tissues. Thus, HAp has attracted the interest of the scientific community in the medicine field, material science, and tissue engineering areas in many years. However, due to the special characteristics of this material, HAp is also studied for various applications including fluorescent lamps, fuel cells, and adsorption of harmful metals, as well as catalysts [1]. TiO2 material was concerned almost 48 years ago. For the first time in 1969, the possibility of solar photoelectrolysis of TiO2 was demonstrated; the powdered TiO2 was studied in the 1980s and TiO2 film photocatalysis

© 2016 The Author(s). Licensee InTech. 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. © 2018 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.

in the 1990s [2]. TiO2 photocatalysis (in particles or films) has gained much attention because the material has high stability, low cost, and nontoxicity and can be easily fabricated by many processes including precipitation, hydrothermal, sol-gel, plasma, etc. The research on TiO2 /hydroxyapatite photocatalytic materials may derive from researches on TiO2 /hydroxyapatite composites in medical industry. To achieve biocompatibility, osteoconduction, and osseointegration, the surface of titanium or its alloy, which is used as a permanent implant material, must be modified by developing hydroxyapatite (HAp, Ca10(PO4 )6 (OH)2 ) coating on the surface [3–9]. By inserting TiO2 inner layer between HAp coating and Ti substrate, the adhesion strength of the coating and the substrate increases [10]. The addition of TiO2 inner layer is expected to reduce a thermal expansion mismatch of the layers, to improve the bonding strength between the HAp layer and Ti substrate, to prevent the corrosion of the Ti substrate, as well as to obtain an abundance of surface hydroxyl and superoxide radical groups, sequentially, to achieve a surface free of cracks and a high adhesion of the modified surface to the substrate [7, 11–13]. The reactions of photocatalysis occur on the surface of TiO2 ; thus, both the surface properties and the mass transfer of the pollutant and degraded products onto the substrate affect to its photocatalytic activity. HAp, which is known as a material having a large surface area and high adsorption ability, may play the role of a transparent or semitransparent to allow UV and visible radiation to pass through it. In fact, there are numerous materials having larger surface area and higher adsorption ability such as silica gel, zeolite, activated carbon, etc. However, besides these general properties, HAp material was investigated as a support for photocatalysis due to the generation of active superoxide anion radicals (O2 ˙ −) under UV irradiation [14]; thus, the TiO2 /HAp photocatalysts have gained the interest of many scientists. In this chapter, the role of hydroxyapatite in TiO2 /HAp photocatalytic materials will be analyzed and evaluated to supply a clear view of the composites.

2–10 nm in thickness. This composite material may adsorb contaminants without exposure to

contaminants such as viruses, bacteria, etc. but also decompose these compounds by the photocatalytic process. Hirakura et al. [22] presented that lysozyme (LSZ) and bovine serum albumin (BSA) can be monomolecularly absorbed on HAp by using two types of fibrous crys-

remove of the specific proteins by the absorbing and decomposing under UV irradiation. Thus, the photocatalytic activity for the decomposition of proteins could be controlled with the adsorption on the surface of the nanostructured HAp crystals. The adsorption of BSA

research explained that the adsorption of the acidic protein BSA occurred at Ca2+ sites of the HAp component which contained a large number of pores supporting to the physical

deposition. The aerosol-deposition films almost fully covered the substrate (glass) and are not porous but extremely rough microstructure. The deposited films maintained good adhesion with the substrate, and the film's pencil hardness was over 9H. The aerosol-deposition TiO2

40 wt% β-TCP composite film has two different phases distinctly: white regions of β-TCP and

formed by the collision of the accelerated particles with high kinetic energy during deposi-

β-TCP adsorbent crystalline phase show the high photocatalytic activity under both UV and

was also evaluated in the report of Ma et al. [25]. This composite membrane, which was fabricated by a facile two-step approach, involves sol–gel process and calcination was a microporous membrane structure with average 0.8 m pore size, which comprised of Ag-TiO2

any structure-directing agent. The result proved the increased photocatalytic activity of the composite results from the combination of adsorption capacity of HAp and the high photo-

ites, especially in atmospheric environment [27–29]. Komazaki et al. [27] collected NO(x) by

Other reports also concerned to the adsorption properties of HAp in the HAp/TiO2

/HAp/Al2

O3

O3

/HAp composite with mosaic structure via a facile route without

/HAp composite was also presented in the publication of Katayama et al. [23]. The

photocatalyst on exposure to light. As a

http://dx.doi.org/10.5772/intechopen.81092

/HAp materials not only absorb organic

anatase can selectively

117

/HAp

compos-

/HAp composites can be used to pure the air or play the role of

Evaluation of the Role of Hydroxyapatite in TiO2/Hydroxyapatite Photocatalytic Materials

nanocrystallites with nanoscaled β-TCP crystallites

photocatalytic crystallites with a dispersion of

bioceramic composite membrane

disk support. HAp component

provided powerful photocata-

and HAp. The research shows

produces reactive oxygen species

light, and the contaminants are decomposed by TiO2

Other literatures show that the photocatalytic TiO2

the antimicrobial and antifungal coating with HAp absorber.

, 5–50 nm-sized TiO2

composite layer with a thickness of 10 m overlaid on α-Al2

acted as a highly efficient bacterial adsorbent, while Ag-TiO2

an annular diffusion scrubber coated with a mixture of TiO2

that HAp plays the role of adsorption material, while TiO2

tion. The films consisting of nano-sized TiO2

dark conditions due to adsorption effect of β-TCP.

The adsorption role of the HAp layer in Ag-TiO2

.

tals elongated in the c-axis. HAp-nanostructured crystals doping TiO2

result, the photocatalytic TiO2

Ryu et al. [24] fabricated the TiO2

lytic attack toward *E. coli* strains.

Xie et al. [26] fabricated TiO2

catalytic activity of TiO2

on TiO2

adsorption.

dark regions of TiO2

#### **2. Hydroxyapatite plays the role of adsorption material in the photocatalytic TiO2 /HAp composites**

Adsorption of biological, organic-chemical molecules on the HAp surface is generally influenced by its physicochemical properties including crystallite size, pore structure, morphology of particles, or coatings [15] which directly depend on the synthesis methods. The synthesis methods of HAp particles include solid-state reaction; sol-gel, plasma, and hydrothermal technique; layer hydrolysis of other calcium phosphate salts; etc. [16–18], while those of HAp coating include sol-gel, chemical vapor deposition, pulsed laser deposition, RF magnetron sputtering, spray pyrolysis, etc. [7, 19, 20]. The choice of a specific method depends on the purpose of research which is synthesis, characteristics, or application of pure HAp or those of TiO2 /HAp composites.

The adsorption property of HAp in the photocatalytic TiO2 /HAp composites has been reported in many researches. Nonami et al. [21] soaked TiO2 powder in a simulated physiological solution containing phosphate ions for periods of about 1 h at 37°C. The apatite film with a thickness of approximately 0.7 μm has formed on the approximately 0.3 μm thick TiO2 layer. The crystals have a plate-like shape, measuring approximately 0.1–0.5 μm in length and 2–10 nm in thickness. This composite material may adsorb contaminants without exposure to light, and the contaminants are decomposed by TiO2 photocatalyst on exposure to light. As a result, the photocatalytic TiO2 /HAp composites can be used to pure the air or play the role of the antimicrobial and antifungal coating with HAp absorber.

in the 1990s [2]. TiO2

116 Photocatalysts - Applications and Attributes

on the surface [3–9]. By inserting TiO2

−) under UV irradiation [14]; thus, the TiO2

of many scientists. In this chapter, the role of hydroxyapatite in TiO2

**/HAp composites**

The adsorption property of HAp in the photocatalytic TiO2

reported in many researches. Nonami et al. [21] soaked TiO2

materials will be analyzed and evaluated to supply a clear view of the composites.

Adsorption of biological, organic-chemical molecules on the HAp surface is generally influenced by its physicochemical properties including crystallite size, pore structure, morphology of particles, or coatings [15] which directly depend on the synthesis methods. The synthesis methods of HAp particles include solid-state reaction; sol-gel, plasma, and hydrothermal technique; layer hydrolysis of other calcium phosphate salts; etc. [16–18], while those of HAp coating include sol-gel, chemical vapor deposition, pulsed laser deposition, RF magnetron sputtering, spray pyrolysis, etc. [7, 19, 20]. The choice of a specific method depends on the purpose of research which is synthesis, characteristics, or application of pure HAp or those of

ological solution containing phosphate ions for periods of about 1 h at 37°C. The apatite film with a thickness of approximately 0.7 μm has formed on the approximately 0.3 μm thick TiO2 layer. The crystals have a plate-like shape, measuring approximately 0.1–0.5 μm in length and

**2. Hydroxyapatite plays the role of adsorption material in the**

TiO2

(O2 ˙

TiO2

**photocatalytic TiO2**

/HAp composites.

photocatalysis (in particles or films) has gained much attention because

/hydroxy-

) coating

inner

; thus,

/HAp photocatalytic

/HAp composites has been

powder in a simulated physi-

)6 (OH)2

inner layer between HAp coating and Ti substrate, the

/HAp photocatalysts have gained the interest

the material has high stability, low cost, and nontoxicity and can be easily fabricated by many processes including precipitation, hydrothermal, sol-gel, plasma, etc. The research on

apatite composites in medical industry. To achieve biocompatibility, osteoconduction, and osseointegration, the surface of titanium or its alloy, which is used as a permanent implant

layer is expected to reduce a thermal expansion mismatch of the layers, to improve the bonding strength between the HAp layer and Ti substrate, to prevent the corrosion of the Ti substrate, as well as to obtain an abundance of surface hydroxyl and superoxide radical groups, sequentially, to achieve a surface free of cracks and a high adhesion of the modified surface

both the surface properties and the mass transfer of the pollutant and degraded products onto the substrate affect to its photocatalytic activity. HAp, which is known as a material having a large surface area and high adsorption ability, may play the role of a transparent or semitransparent to allow UV and visible radiation to pass through it. In fact, there are numerous materials having larger surface area and higher adsorption ability such as silica gel, zeolite, activated carbon, etc. However, besides these general properties, HAp material was investigated as a support for photocatalysis due to the generation of active superoxide anion radicals

/hydroxyapatite photocatalytic materials may derive from researches on TiO2

adhesion strength of the coating and the substrate increases [10]. The addition of TiO2

to the substrate [7, 11–13]. The reactions of photocatalysis occur on the surface of TiO2

material, must be modified by developing hydroxyapatite (HAp, Ca10(PO4

Other literatures show that the photocatalytic TiO2 /HAp materials not only absorb organic contaminants such as viruses, bacteria, etc. but also decompose these compounds by the photocatalytic process. Hirakura et al. [22] presented that lysozyme (LSZ) and bovine serum albumin (BSA) can be monomolecularly absorbed on HAp by using two types of fibrous crystals elongated in the c-axis. HAp-nanostructured crystals doping TiO2 anatase can selectively remove of the specific proteins by the absorbing and decomposing under UV irradiation. Thus, the photocatalytic activity for the decomposition of proteins could be controlled with the adsorption on the surface of the nanostructured HAp crystals. The adsorption of BSA on TiO2 /HAp composite was also presented in the publication of Katayama et al. [23]. The research explained that the adsorption of the acidic protein BSA occurred at Ca2+ sites of the HAp component which contained a large number of pores supporting to the physical adsorption.

Ryu et al. [24] fabricated the TiO2 -β-TCP nanocomposite photocatalytic thin films by aerosol deposition. The aerosol-deposition films almost fully covered the substrate (glass) and are not porous but extremely rough microstructure. The deposited films maintained good adhesion with the substrate, and the film's pencil hardness was over 9H. The aerosol-deposition TiO2 - 40 wt% β-TCP composite film has two different phases distinctly: white regions of β-TCP and dark regions of TiO2 , 5–50 nm-sized TiO2 nanocrystallites with nanoscaled β-TCP crystallites formed by the collision of the accelerated particles with high kinetic energy during deposition. The films consisting of nano-sized TiO2 photocatalytic crystallites with a dispersion of β-TCP adsorbent crystalline phase show the high photocatalytic activity under both UV and dark conditions due to adsorption effect of β-TCP.

The adsorption role of the HAp layer in Ag-TiO2 /HAp/Al2 O3 bioceramic composite membrane was also evaluated in the report of Ma et al. [25]. This composite membrane, which was fabricated by a facile two-step approach, involves sol–gel process and calcination was a microporous membrane structure with average 0.8 m pore size, which comprised of Ag-TiO2 /HAp composite layer with a thickness of 10 m overlaid on α-Al2 O3 disk support. HAp component acted as a highly efficient bacterial adsorbent, while Ag-TiO2 provided powerful photocatalytic attack toward *E. coli* strains.

Xie et al. [26] fabricated TiO2 /HAp composite with mosaic structure via a facile route without any structure-directing agent. The result proved the increased photocatalytic activity of the composite results from the combination of adsorption capacity of HAp and the high photocatalytic activity of TiO2 .

Other reports also concerned to the adsorption properties of HAp in the HAp/TiO2 composites, especially in atmospheric environment [27–29]. Komazaki et al. [27] collected NO(x) by an annular diffusion scrubber coated with a mixture of TiO2 and HAp. The research shows that HAp plays the role of adsorption material, while TiO2 produces reactive oxygen species under ultraviolet light (UV) illumination such as super oxide (O2 −), hydroxyl radical (OH°), and peroxyhydroxyl radical (HO2 °), by which nitric oxide is oxidized to nitrogen dioxide and is further oxidized to nitric acid.

Ozeki et al. [28] fabricated TiO2 /HAp thin film by sputtering on glass using a radio-frequency magnetron. The film has a higher decomposition rate of formaldehyde gas than either the TiO2 or the HAp film alone. However, in the bacterial survival test, survival of cells on the TiO2 /HAp thin film is higher than that on the TiO2 film, which indicates that the TiO2 /HAp thin film has a lower bactericidal effect than the pure TiO2 film.

Liu et al. [29] synthesized HAp-modified N-TiO2 by a facile wet-chemical method and evaluated its photocatalytic activity by the decomposition of gaseous acetone under visible-light irradiation. The results demonstrated that 10%-HAp-N-TiO2 sample shows the best photocatalytic activity and the remarkable photocatalytic activity may arise from the synergism between adsorption on HAp and photoactivity by TiO2 which generated oxygen-reactive species diffusing and reacting with the molecules located on the HAp.

Masato Wakamura et al. have many publications on the field [30–32]. The authors evaluated the adsorption behaviors of proteins onto photocatalytic Ti4+-doped calcium hydroxyapatite (TiHAp) particles which was synthesized by the coprecipitation method using Ca(NO3 )2 , Ti(SO4 )2 , H3 PO4 , and NH4 OH original materials [31]. The result showed that all the adsorption isotherms of bovine serum albumin (BSA), myoglobin (MGB), and lysozyme (LSZ) from 1 × 10−4 mol/dm3 KCl solution were Langmuir type. Generally, the saturated amounts of adsorbed protein value on the TiHAp were much higher than CaHAp.

vacancies formed on the surface of the excited PO4

**Scheme 1.** Photocatalytic mechanism of cyanide degradation by Pd-TiO2

hydroxyapatite nanoparticle was higher than those of Pd-TiO2

the formation of O2

TiO2

apatite in the photocatalytic TiO2

erated electrons and holes in TiO2

enhancing quantum efficiency of photocatalysts.

efficiency of organic pollutants also increases.

The promoted impact of hydroxyapatite in the photocatalytic TiO2

sented in the report of Liu et al. [37]. The photocatalytic activities of Ag-TiO2

3− group in visible illumination will lead to

http://dx.doi.org/10.5772/intechopen.81092

and TiO2

nanoparticles is promoted by the internal polarization of

119

nanoparticles. The

/HAp composites was pre-

nanopar-

˙− and will attack the surrounding organic molecules adsorbed on HAp.

Evaluation of the Role of Hydroxyapatite in TiO2/Hydroxyapatite Photocatalytic Materials

The research [34] investigated that the amount of hydroxyl radials formed in the Pd-TiO2

absorption of radiation in the visible-light region, the small recombination rate of the electronhole pair, as well as the high surface area of HAp material are the promoted impacts of hydroxy-

The recombination rate of the electron-hole pair is also concerned in another publication [35].

ticles which was supported on electrically polarized HAp films. The separation of photogen-

the HAp support, and consequently, the recombination of charge carriers is mitigated. It can be concluded that the materials with large internal polarization can be used in strategies for

The surface area of HAp material was concerned in the research of Kobayashi et al. [36]. The

synthesized by a facile wet-chemical strategy were evaluated by photocatalytic oxidation decomposition of acetone in air under visible-light illumination. The high photocatalytic activity

/HAp photocatalytic coating was deposited by gas tunnel-type plasma spraying using powders with nano-sized grains. The photocatalytic reaction occurs on powder surface; as the grain size becomes smaller, the specific surface area becomes larger, and the degradation

/HAp composites.

Zhang et al. [35] demonstrated the increased photocatalytic performance of TiO2

#### **3. The promoted impact of hydroxyapatite in the photocatalytic TiO2 /HAp composites**

Beside the role as an adsorption material in the photocatalytic TiO2 /HAp composites, some recent researches have investigated that hydroxyapatite promotes photocatalytic degradation because it can play the role of a support for photocatalysts. The activity of HAp is caused by the generation of active superoxide anion radicals (O2 ˙−) due to a change in the electronic state of the surface PO4 3− group under UV irradiation [14].

Hu et al. [33] proved the promotion role of HAp to TiO2 in photocatalytic degradation by using experimental observations and kinetic modeling. The derived kinetic parameters including reaction rate constant, Langmuir adsorption constant, apparent activation energy, etc. confirm that the activity of TiO2 /HAp composite is more effective than that of TiO2 . A negative effect of HAp on the photo absorption ability of the TiO2 /HAp composite is characterized by UV-vis reflectance spectra. The existence of superior chemisorption between HAp and the organic molecules leads to a better performance of TiO2 /HAp for photocatalytic degradation.

Mohamed et al. [34] determined photocatalytic activities of Pd-TiO2 -hydroxyapatite nanoparticles, which are synthesized by a template-ultrasonic-assisted method, by photocatalytic removal of cyanide under visible-light irradiation (**Scheme 1**). **Scheme 1** shows that the

**Scheme 1.** Photocatalytic mechanism of cyanide degradation by Pd-TiO2 -HAp [34].

under ultraviolet light (UV) illumination such as super oxide (O2

/HAp thin film is higher than that on the TiO2

Liu et al. [29] synthesized HAp-modified N-TiO2

, and NH4

**/HAp composites**

thin film has a lower bactericidal effect than the pure TiO2

irradiation. The results demonstrated that 10%-HAp-N-TiO2

cies diffusing and reacting with the molecules located on the HAp.

adsorbed protein value on the TiHAp were much higher than CaHAp.

Beside the role as an adsorption material in the photocatalytic TiO2

3− group under UV irradiation [14].

the generation of active superoxide anion radicals (O2

Hu et al. [33] proved the promotion role of HAp to TiO2

effect of HAp on the photo absorption ability of the TiO2

organic molecules leads to a better performance of TiO2

Mohamed et al. [34] determined photocatalytic activities of Pd-TiO2

**3. The promoted impact of hydroxyapatite in the photocatalytic**

between adsorption on HAp and photoactivity by TiO2

and peroxyhydroxyl radical (HO2

is further oxidized to nitric acid.

118 Photocatalysts - Applications and Attributes

Ozeki et al. [28] fabricated TiO2

TiO2

TiO2

Ti(SO4 )2 , H3 PO4

**TiO2**

1 × 10−4 mol/dm3

of the surface PO4

firm that the activity of TiO2

−), hydroxyl radical (OH°),

/HAp

)2 ,

°), by which nitric oxide is oxidized to nitrogen dioxide and

/HAp thin film by sputtering on glass using a radio-frequency

film.

OH original materials [31]. The result showed that all the adsorp-

KCl solution were Langmuir type. Generally, the saturated amounts of

film, which indicates that the TiO2

by a facile wet-chemical method and evalu-

sample shows the best photo-

/HAp composites, some

. A negative

˙−) due to a change in the electronic state

in photocatalytic degradation by using

/HAp composite is characterized by

/HAp for photocatalytic degradation.

which generated oxygen-reactive spe-

magnetron. The film has a higher decomposition rate of formaldehyde gas than either the

ated its photocatalytic activity by the decomposition of gaseous acetone under visible-light

catalytic activity and the remarkable photocatalytic activity may arise from the synergism

Masato Wakamura et al. have many publications on the field [30–32]. The authors evaluated the adsorption behaviors of proteins onto photocatalytic Ti4+-doped calcium hydroxyapatite (TiHAp) particles which was synthesized by the coprecipitation method using Ca(NO3

tion isotherms of bovine serum albumin (BSA), myoglobin (MGB), and lysozyme (LSZ) from

recent researches have investigated that hydroxyapatite promotes photocatalytic degradation because it can play the role of a support for photocatalysts. The activity of HAp is caused by

experimental observations and kinetic modeling. The derived kinetic parameters including reaction rate constant, Langmuir adsorption constant, apparent activation energy, etc. con-

UV-vis reflectance spectra. The existence of superior chemisorption between HAp and the

ticles, which are synthesized by a template-ultrasonic-assisted method, by photocatalytic removal of cyanide under visible-light irradiation (**Scheme 1**). **Scheme 1** shows that the

/HAp composite is more effective than that of TiO2

or the HAp film alone. However, in the bacterial survival test, survival of cells on the

vacancies formed on the surface of the excited PO4 3− group in visible illumination will lead to the formation of O2 ˙− and will attack the surrounding organic molecules adsorbed on HAp.

The research [34] investigated that the amount of hydroxyl radials formed in the Pd-TiO2 hydroxyapatite nanoparticle was higher than those of Pd-TiO2 and TiO2 nanoparticles. The absorption of radiation in the visible-light region, the small recombination rate of the electronhole pair, as well as the high surface area of HAp material are the promoted impacts of hydroxyapatite in the photocatalytic TiO2 /HAp composites.

The recombination rate of the electron-hole pair is also concerned in another publication [35]. Zhang et al. [35] demonstrated the increased photocatalytic performance of TiO2 nanoparticles which was supported on electrically polarized HAp films. The separation of photogenerated electrons and holes in TiO2 nanoparticles is promoted by the internal polarization of the HAp support, and consequently, the recombination of charge carriers is mitigated. It can be concluded that the materials with large internal polarization can be used in strategies for enhancing quantum efficiency of photocatalysts.

The surface area of HAp material was concerned in the research of Kobayashi et al. [36]. The TiO2 /HAp photocatalytic coating was deposited by gas tunnel-type plasma spraying using powders with nano-sized grains. The photocatalytic reaction occurs on powder surface; as the grain size becomes smaller, the specific surface area becomes larger, and the degradation efficiency of organic pollutants also increases.

The promoted impact of hydroxyapatite in the photocatalytic TiO2 /HAp composites was presented in the report of Liu et al. [37]. The photocatalytic activities of Ag-TiO2 -HAp powders synthesized by a facile wet-chemical strategy were evaluated by photocatalytic oxidation decomposition of acetone in air under visible-light illumination. The high photocatalytic activity of the Ag-TiO2 -HAP hybrids could be attributed to its strong absorption in the visible-light region, low recombination rate of the electron-hole pair, and large BET-specific surface area.

levels corresponded with the characterization of titanium (IV) in TiO2

states of titanium in the annealing process of TiO(OH)2

/HAp material calculated by the DRS measurement was 3.6 eV, while those

may affect the absorption ability of radiation in the visible-light region; consequently, in some cases, it is a demoted impact which reduces the photocatalytic activity of TiO2

nations which is generally influenced by the physicochemical properties of HAp material. On the other hand, the promoted impacts of hydroxyapatite in the composites are experimentally collected in the specific cases including the absorption of radiation in the visible-light region, the small recombination rate of the electron-hole pair, and the high surface area of HAp com-

continuously researched and evaluated to obtain the best composite in the photocatalytic

This work is supported by Basic Science Research Programs (No. 2015R1D1A1A01056983) through the NRF, Korea, funded by the Ministry of Education and (No. 2018R1A2B6002194)

1 Department of Chemistry, Ho Chi Minh City University of Education, Ho Chi Minh City,

2 Department of Materials Science and Engineering, Korea Advanced Institute of Science

[1] Fihri A, Len C, Varma RS, Solhy A. Hydroxyapatite: A review of syntheses, structure and applications in heterogeneous catalysis. Coordination Chemistry Reviews. 2017;

and Kwangsoo No2

\*, Seungbum Hong2

\*Address all correspondence to<EMAIL_ADDRESS>
were 5.3 and 3.2 eV, respectively. The shifts may relate oxidation

Evaluation of the Role of Hydroxyapatite in TiO2/Hydroxyapatite Photocatalytic Materials

/HAp composites, HAp plays the role of an adsorption of contami-

/hydroxyapatite photocatalytic materials has been

bandgap of TiO2

composites.

**4. Conclusion**

In the photocatalytic TiO2

**Acknowledgements**

**Author details**

Vietnam

**References**

**347**:48-76

Linh Nguyen Thi Truc1

ponent. However, the role of HAp in TiO2

field, as well as in other applications.

by the Ministry of Science and ICT.

and Technology, Daejeon, South Korea

of pure HAp and TiO2

. The experimental

/HAp

121

/HAp gel. The increase of bandgap

http://dx.doi.org/10.5772/intechopen.81092

Aramendia et al. [38] studied on the TiO2 /natural phosphate material synthesized by sol-gel process and tested the photocatalytic activity in the photo-oxidation process of propan-2-ol. The results show that the presence of natural phosphate makes a retardation of TiO2 crystallization and creates a certain interaction with titanium.

Nishikawa et al. [39] indicated that the difference in photocatalytic activity of the materials containing HAp and without HAp may be originated from the properties of the electronic state that forms the valence band of HAp-containing PO4 group.

Li et al. [40] synthesized Ti-substituted hydroxyapatite (TiHAp) by the coprecipitation method. The adsorption and photocatalytic degradation of bisphenol A (BPA, an environmental endocrine disrupting chemical) over TiHAp and P25 TiO2 photocatalysts were studied using liquid chromatography-mass spectrometry. The results indicated that the adsorption of BPA on TiHAp and TiO2 obeyed the Langmuir adsorption equation, and the adsorption capacity and photocatalytic degradation activity of BPA of TiHAp material were much higher than those of TiO2 . To explain for the results, the authors presented that the zeta potentials of TiHAp, and TiO2 did not show significant difference when the pH was about seven, suggesting that the surface charge was not the reason for the different adsorption capacities of the particles. In addition, the specific surface area and average pore diameter of TiHAp and TiO2 were comparable, so these would not lead to the different adsorption capacities either. In fact, TiHAp material is produced from a substitution of some Ca sites in HAp by Ti, which resulted in multiple Ti-OH groups on the TiHAp surface. Large amounts of phosphates and hydroxyls in the crystal lattice of TiHAp can adsorb the hydroxyls of BPA by hydrogen bonding.

The bandgap of TiO2 /HAp materials were mentioned in some publications because it decides the energy separation between valence and conduction bands, the quantum effect, as well as the effect of visible-light utilization, etc. The bandgap of HAp was reported in [41] to be 3.95 eV by photoluminescence measurement; meanwhile, in other reports, the bandgap of HAp was calculated to be around 4.51–5.4 eV [42, 43]. The bandgap of TiO2 /HAp composites calculated by UV-vis diffuse reflectance spectra was between 3.06 and 3.08 eV while that of pure TiO2 was broader (3.12 eV) [44]. However, the bandgap of Ti-substituted hydroxyapatite evaluated by both experimental and theoretical methods was 3.65 eV [45]. Masato Wakamura et al. determined the effect of Ti substitution in HAp synthesized by the coprecipitation method on the bandgap which was compared with that of a typical anatase-TiO2 photocatalytic powder. The experimentally obtained optical bandgap energies of TiHAp, HAp, and TiO2 powder measured by diffuse reflectance spectroscopy were 3.65, >6, and 3.27 eV, respectively. The authors explained the increase of bandgap in TiHAp due to hybridizing of Ti 3d orbital with O 2p orbital and forming an internal state in the HAp bandgap, consequently, causing the absorption-edge lowering of TiHAp. Researching on the acetaldehyde gas decomposition of TiHAp by UV with VIS irradiation, the authors investigated the increase of activity comparing with when UV irradiation alone was used. Linh et al. [46] also indicated the shift of bandgap of TiO2 /HAp material synthesized by the hydrothermal process containing TiO(OH)2 and HAp gels. The binding energy values of Ca 2p, P 2p, and O1s levels are related to hydroxyapatite phase, whereas those of Ti 2p levels corresponded with the characterization of titanium (IV) in TiO2 . The experimental bandgap of TiO2 /HAp material calculated by the DRS measurement was 3.6 eV, while those of pure HAp and TiO2 were 5.3 and 3.2 eV, respectively. The shifts may relate oxidation states of titanium in the annealing process of TiO(OH)2 /HAp gel. The increase of bandgap may affect the absorption ability of radiation in the visible-light region; consequently, in some cases, it is a demoted impact which reduces the photocatalytic activity of TiO2 /HAp composites.