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Abstract

The high demand for quality in the manufacture of absorbent hygiene products requires the adhesive bonds between layers to be as uniform as possible. An experimental study was conducted on two industrial multihole melt blowing nozzle designs used for hot-melt adhesive applications for hygiene products, in order to study two defects that influence the quality of the adhesive bond: fibre breakup, resulting in contamination, and the presence of shots, undesirable lumps that end up in the finished product. To this end, the fibre dynamics were captured at the nozzle exit region by using high-speed imaging. From the results it was observed that die drool is the main source of shot formation, while fibre breakup occurs as a result of applying a sufficiently large force in the direction perpendicular to the fibre. In addition, three dimensionless parameters were defined, the first two being the air-polymer flux ratio and the dimensionless temperature ratio, both of which represent the operating conditions, and the remaining one being the force ratio, which represents the nozzle geometry. The effect of these parameters on fibre breakup and shot formation was studied and the results indicate that both the operating conditions and the nozzle geometry were responsible for the onset of the fibre breakup and for the formation of shots. More precisely, both defects turned out to be dominated by the air-polymer flux ratio and the air tilt angle. The results that emerge from this study are useful for the enhancement of industrial melt blowing nozzles.

Keywords

Melt blown shot cohesive breakup die drool melt blowing regions

Introduction

In the melt blowing (MB) process, a polymer stream is molten and extruded through a set of nozzles and it is immediately blown and attenuated by a set of convergent high-velocity hot-air jets. Among all the industrial processes available for producing nonwoven microfibers, the MB process is particularly effective because it allows webs with a high surface area per unit volume to be produced, usually without requiring any finishing operations [1]. More recently, the MB process has begun to be used to produce hot-melt adhesives, which are used in various industrial processes to bond different kinds of products, such as products that include nonwoven fabrics. Particularly, in melt blowing adhesives, the final fibre diameters range from 10 to 20 µm [2]. The MB system comprises a MB head that supports a set of nozzles, which can have different designs [3], depending primarily on the desired characteristics of the application pattern. A scheme representing the principal components in the melt blowing process is shown in Figure 1. However, because each nozzle design leads to different defects on the produced application pattern, the aim of this article is to study two MB nozzle designs and explain how the geometric and operational factors of each design lead to their respective defects on the adhesive patterns. Depending on production requirements, the fibres can be extruded either continuously or discontinuously, but in the present work only the continuous extrusion of polymer by non–contact nozzle designs are studied. Non-contact nozzles present some advantages when compared with contact nozzles, for example they present a better clog resistance and permit bonding uneven surfaces. They are also better for bonding temperature sensitive materials.

Figure 1.

Sketch of the melt blowing process.

Two aspects of the MB process are not yet fully understood, namely the fibre breakup process and shot formation (or drop formation). The first is a key subject because fibre breakup reduces the fibre’s mean diameter, thus increasing the fibre’s surface area and the uniformity of the adhesive bond. However, excessive fibre breakup causes undesirable defects, such as flies. To date, little information has been published about the factors that lead to fibre breakup in the MB process. C. Ellison et al. [4] observed that, for viscoelastic fluids, fibre breakup depends on the polymer’s melt viscosity and elasticity, which control the extensional stress, the operating temperature and the air-polymer flux ratio. An increase either in the temperature or in the air-polymer flow ratio increases the fibre breakup. Moreover, they associated the onset of this phenomenon with the presence of spherical particles dispersed amongst the fibres. Y. Wang et al. [5] conducted uniaxial stretching experiments on styrene-butadiene rubbers to study fibre breakup of polymer melts. The authors reported that the yield extensional strain $(ε_{y})$ of the polymer increases with both increasing extensional rates $(\dot{ε})$ and relaxation times $(τ_{R})$ , for which they obtained $ε_{y} \sim {(\dot{ε} τ_{R})}^{1 / 3}$ . Parameter $τ_{R}$ represents the Rouse relaxation time of the polymer, which is proportional to the polymer’s squared molecular weight [6]. A similar behaviour was observed on shear experiments, for which they have obtained yield shear strain $γ_{y} \sim {(\dot{γ} τ_{R})}^{1 / 3}$ . More recently, Wanli Han et al. [7] carried out a theoretical study of the breakup of MB nanofibers. In line with the work of Ellison et al., they also observed the presence of small droplets within the melt blown web and stated that these are indeed related to the fibre breakup but do not represent shots, which are usually described as “clumps of polymer that appear as a result of the operating conditions” [1]. They concluded that fibre diameter, operating temperature and air pressure are all important factors that affect fibre breakup.

Regarding shot formation, it is known that in hot-melt adhesive applications this is a matter of special importance because the presence of shots in the application pattern reduces the uniformity of the adhesive bond and therefore the quality of the final product, as well as increases the amount of adhesive wasted unnecessarily. Shots are also undesirable because they can cause heat distortion if the material is temperature sensitive. The formation of shots is primarily attributed to operating conditions and equipment cleanliness. For instance, R. Breese and Z. Yan devoted two articles to the study of the shot phenomenon [8,9]. In the first one, they studied how the number of shots vary as a function of the die-to-collector distance (DCD) and the web basis weight, and they explained the shot formation process on the basis of three different models: the fibre transformation model, the direct shot formation model and the shot precursor model. However, they concluded that no model alone could explain all the experimental observations and stated that shot formation probably involves all three mechanisms. In the present study, only the direct shot formation model is considered in order to explain the formation of shots. In the same article, Breese and Yan also stated that the shot particle’s diameters are in the range 100–300 µm, a noticeably larger size than the average fibre diameter. In their second article, Breese and Yan presented a computer software tool they developed to analyse MB web images automatically and calculate several characteristics of the shot, such as: the number of shots, the size-distribution, the mean size and the cover rate. What is more, R. Shambaugh describes the fibre shots as undesirable lumps that are larger than 300 µm in diameter while stating that the shots typically form under certain operating conditions [10]. Other authors have affirmed that shot formation is a consequence of the fibre breakup and explained that shot formation is connected to the fibre elastic snap back effect and to the fusion of filaments on the way to the collector screen. In view of this, it can be stated that there is no universal agreement among researchers on the sources of shot formation, and in the present work an additional explanation is given.

The aim of this article is to extend the present knowledge of these two defects by clarifying how the nozzle geometry and the operating conditions of multihole MB nozzle designs can affect both fibre breakup and shot formation. Regarding the first phenomenon, the results obtained in the present work suggest that fibre breakup is affected by both the operating conditions and the nozzle geometry. Achieving fibre breakup conditions would require working with high air and low polymer flow rates and with high polymer temperatures, while using a nozzle design with highly tilted air ducts. Regarding the second phenomenon, the findings suggest that shots form as a result of the accumulation of adhesive at the nozzle tip (die drool), and working at low polymer temperatures and air flow rates and with high polymer flow rates will help to minimize the formation of shots. The present work also demonstrates a lower formation of shots in nozzle designs with highly tilted air ducts.

Experimental

Description of the nozzle geometry

Since the multihole (or laminated) nozzle geometry ensures consistent bonding between adjacent layers and the application pattern produced by contiguous nozzles does not overlap, it is popular among the MB nozzle designs when it comes to the disposal of hot melt in hygiene products. In the present study, two multihole melt blown nozzle designs used to apply hot-melt adhesive to hygiene products intended to absorb bodily fluids were studied, and both designs are industrial prototypes, which were designed based on two existing commercial designs. Laminated melt blown nozzles consists of an assembly of several metallic plates, where the air orifices and the adhesive capillaries lie. The set of plates is screwed to the body of the nozzle. From now on, the two nozzle designs will be named as A Design and B Design respectively. Both nozzle designs were studied on a modular MB head with interchangeable modular units placed side by side. The melt blown head used for the experiments support 5 modules, each holding a nozzle with a given number of polymer exits, with each exit being surrounded by a certain number of air exits $(N_{a / f})$ . Whereas the nosepiece of the A design is of the outset-sharp type, the one of the B design is of the flush-blunt type. The A nozzle design has 5 adhesive capillaries, each of them surrounded by 2 air exits, while the B nozzle design has 12 adhesive capillaries, each of them surrounded by 4 air exits. The A and B nozzle designs have a width of 25 mm. The air and the polymer exit areas will be described in terms of their hydraulic diameters, where the hydraulic diameter of one air exit $(L_{R})$ in the A design is 520 µm, the one in the B nozzle design is 250 µm. The hydraulic diameter of one adhesive capillary $(L_{R f})$ in the A nozzle design is 520 µm, while the one in the B nozzle design is 380 µm. The distance between two consecutive adhesive capillaries $(p)$ in the A design is $9.6 L_{R}$ , while the one in the B design is $8.0 L_{R}$ . The length of the adhesive capillary $(L)$ in the A design is $8.6 L_{R}$ and the one in the B nozzle design is $13.5 L_{R}$ .

In the multihole nozzle designs, the air exits are often non-parallel to the adhesive exits, hence an angle forms between them. The tilt angles between the air exits and the adhesive capillary in the $X Z$ and $Y Z$ planes are different in both designs. The tilt angle in the $X Z$ plane ( $α$ ) is 0 in the A design and is between 30 and 50 degrees in the B nozzle design, while the tilt angle in the $Y Z$ plane ( $β$ ) is non-zero in both designs, being about 2.5 times larger in the A nozzle design than in the B nozzle design. Figure 2 illustrates the geometry of the A and B nozzle designs by depicting one polymer exit, with its surrounding air exits, and it is referred to as a set of exits of the nozzle.

Figure 2.

Dimensions and parameters defining the nozzle designs: (a) A design and (b) B design. Photos of the nozzles dispensing adhesive (c) A design and (d) B design.

For convenience, a unique parameter that indicates the angle between the air ducts and the polymer capillaries will be defined as a function of the previously defined angles $α$ and $β$ and will be referred to as the global tilt angle ( $ψ$ ). The global tilt angle is defined as the angle between the air duct and the adhesive capillary in the radial direction of the nozzle, in the plane defined by the air duct axis and the polymer capillary axis. For convenience, Figure 3 represents a quarter from a general multihole nozzle exit, in which the polymer exit is surrounded by a total of four air exits. In order to give the proper mathematical definition of the global tilt angle, the unit vector collinear with the kth air duct ( ${\hat{r}}_{k}$ ) will be expressed as a function of $α$ and $β$ , as shown in equation (1) (see Figures 2 and 3):

{\hat{r}}_{k} = \frac{1}{\sqrt{1 + {(\sin α \sin β)}^{2}}} (\cos β \sin α \hat{X} + \sin β \cos α \hat{Y} + \cos β \cos α \hat{Z})

(1)

Figure 3.

Representation of one quarter of a general multihole nozzle.

Where $\hat{X}$ , $\hat{Y}$ , and $\hat{Z}$ are the unit vectors that point in the directions of $X$ , $Y$ and $Z$ , respectively. Then, the global tilt angle, which is defined as the angle between the air duct ( ${\hat{r}}_{k}$ ) and the polymer capillary, which is collinear with the vertical unit vector $\hat{Z}$ , can be calculated by performing the dot product between both unit vectors, which results in equation (2):

\cos ψ = \frac{\cos β \cos α}{\sqrt{1 + {(\sin α \sin β)}^{2}}}

(2)

From equation (2), it can be verified that if $β = 0$ , then $ψ = α$ and vice versa. If we substitute the values of $α$ and $β$ for the nozzles involved in the present work and proceed to calculate $ψ$ by using equation (2), the angle $ψ$ of the B nozzle design turns out to be 5.6 times larger than the one of the A nozzle design. This ratio will turn out to be especially important when analysing fibre breakup in Fibre breakup section.

In addition to the global tilt angle ( $ψ$ ), a second angle will be defined between the projection of the air duct axis onto the $X Y$ plane and the $Y$ -axis. This angle will be called as the precession angle and represented by the Greek letter $θ$ (see Figure 3). Unlike the angle $ψ$ , this angle measures how close the air exits are placed to each other in the $X$ and $Y$ directions. In the A Design the air and polymer exits are aligned in the $Y$ direction, then angle $θ$ is zero. However, in the B Design angle $θ$ is not zero. As Figure 3 depicts, the origin of the coordinate system is located at the centre of the polymer exit.

Polymer characterization

Hot melt adhesives are mainly composed of a high molecular copolymer, resin and wax [11]. This type of adhesive is commonly solid at ambient temperature, becomes tacky as it is molten and forms a strong adhesive bond as it cools. The adhesive used in the present study is a typical commercial construction polymer (whose commercial name is PM357V) used in the manufacturing of hygiene products, such as diapers and sanitary napkins. The adhesive has been manufactured by company Savaré. More precisely, this is a styrene butadiene styrene (SBS) based adhesive, whose formula includes other components such as hydrocarbon resins, oil minerals and less than 3% in other additives. Based on thermoplastic rubber, its viscosity is 1.6 and 7.3 Pa.s at 170°C and 130°C, respectively, presenting a similar thermal behaviour to other hot-melt adhesives used for sanitary applications [12]. Its softening point is about 76°C (ASTM E 28, water), its flash point is over 200°C and its maximum running temperature is 170°C, which establishes the maximum value for the polymer operation temperature on the experiments.

Figure 4 shows the shear viscosity curve of the adhesive, which is given by the Cross Model, from which $µ_{o}$ is the zero-shear viscosity, $λ$ is the relaxation time and $α$ an exponent that accounts for the adhesive’s non-linearity with shear rate. The viscosity measurements have been performed using a parallel plate rheometer, for three different distances between plates (gap width), namely 300, 500 and 700 µm.

Figure 4.

Adhesive’s viscosity as the function of the shear rate at $150$ °C.

Experimental facility and recording parameters

Given the fast fibre dynamics present in the melt blowing of hot-melt adhesives, a high-speed camera must be used to capture the fibre shot and the fibre breakup in the nozzle exit region. Figure 5 shows an image and a sketch of the experimental setup that was used for this commitment. The experimental facility comprises the MB head with the set of nozzles, supported by a 3-degree-of-freedom manual positioning mechanism, the high-speed camera (Photron FASTCAM Mini AX200) and a cold light source (Tungsten Light Head COOLH dedocool). In order to record the images under backlight conditions, the camera was placed facing the light source. To adequately focus on the fibre, the camera was mounted on a micrometric platform, which allowed very precise positioning along the $X$ direction. In addition, the platform was mounted onto a tripod that permitted the positioning of the camera along the $Z$ direction. In order to obtain a close view of the adhesive fibre at the central polymer exit, a long-distance microscope (LDM) (INFINITY K2/S) had to be used as the camerás objective. Considering a working distance of 305 mm between the camera and the fibre, the LDM established a maximum field of view (FOV) of about 8.7 × 8.7 mm². By using a level, the camera was positioned fully horizontally, while its vertical position was adjusted by moving the tripod until the top of the cameras FOV matched up with the nozzle exit, within an uncertainty ranging from 60 to 100 µm.

Figure 5.

Experimental facility used for the high-speed imaging: (a) photograph and (b) sketch.

After conducting preliminary tests, the appropriate number of frames per second (fps) required for capturing the fibre dynamics turned out to be 14,400 and 19,200 for the A and B designs, respectively. The resulting FOV for each case comprised 384 × 1024 and 256 × 1024 pix², which translated into 3.3 × 8.7 and 2.2 × 8.7 mm², respectively. A total of 950 frames where recorded in each experiment, which is equivalent to about 65 and 50 ms of recording time for the A and B designs, respectively. The shutter speed was set to 1 µm, which resulted in a proper exposure time that permitted the recording of bright and sharp images, with no blur.

Operating conditions

Three are the operating factors that are usually considered when studying the MB process: the air flow rate $(Q_{a})$ , the adhesive mass flow rate ( ${\dot{m}}_{f}$ ) and the adhesives temperature of operation $(T_{f})$ (i.e. extrusion temperature), where the latter refers to the temperature of the adhesive before extrusion. Based on the literature review, the following typical operating conditions are found in most hot melt adhesive applications (2):

polymer temperature between 138 and 163 degrees Celsius.

polymer mass flow rate is between 0.1 and 10 g/min/hole.

air flow rate between 2.8 and 56.6 L/min/inch.

Table 1 presents the experimental conditions that were set for all the experiments conducted, which vary within the typical range of variation presented above. All these conditions were taken from real industrial hot-melt adhesive applications for hygiene products. The air temperature $(T_{a})$ was not considered as an additional factor and it was kept constant at 200°C on all the experiments, with a precision of ±3°C. A two-level design of experiments that included the three factors considered above was carried out. Additionally, a central point was included in the experimental schedule, totalling 9 experiments on each nozzle design, as outlined in Table 1. The values of $Q_{a}$ and ${\dot{m}}_{f}$ presented in Table 1 correspond both to each polymer exit.

Table 1.

Operating conditions established for performing the HSI.

A design				B Design
#Exp.	$Q_{a}$ [NL/min]	${\dot{m}}_{f}$ [g/min]	$T_{f}$ [^oC]	#Exp.	$Q_{a}$ [NL/min]	${\dot{m}}_{f}$ [g/min]	$T_{f}$ [^oC]
A1	2.6	1.5	140	B1	1.8	0.6	140
A2	5.4	1.5	140	B2	3.7	0.6	140
A3	2.6	3.0	140	B3	1.8	1.3	140
A4	2.6	1.5	170	B4	1.8	0.6	170
A5	5.4	3.0	140	B5	3.7	1.3	140
A6	5.4	1.5	170	B6	3.7	0.6	170
A7	2.6	3.0	170	B7	1.8	1.3	170
A8	5.4	3.0	170	B8	3.7	1.3	170
A9	4.0	2.5	155	B9	2.7	1.0	155

The polymer mass flow rate and temperature were controlled from the MB control system. The adhesive was heated while stored inside the MB tank and also inside the MB head. The adhesive pump provided a volume of 2.5 cubic centimetres of adhesive per each revolution. The precision of the adhesive heating system is about ±0.5°C, as indicated by the manufacturer. At the same time, the airflow rate was controlled with a pressure regulating valve followed by a flow meter with a precision of ∓ (3% o.m.v + 0.3% of the full scale -600 NL/min-), and in the conducted experiments it represented 4 to 6% of the flowrate. The air temperature was controlled from the MB control system, where the air was heated only at the nozzle head, before reaching the nozzles. Before each recording was done, the system was allowed enough time to reach steady state conditions. As is regularly done in continuous hot-melt applications, the polymer mass flow rate and the airflow were both kept constant during the experiments.

In what follows, some dimensionless operating parameters will be defined in order to express the operating conditions of Table 1 in a more compact way. First, a number that represents the ratio between the air and the polymer flow rates will be defined as:

m^{'} = \frac{{\dot{m}}_{a}}{{\dot{m}}_{f}} = \frac{ρ_{o} Q_{a}}{\frac{ρ_{f} v_{c c} n_{rpm}}{N_{f}}}

(3)

This number will be called the air-polymer flux ratio, and it has been used in the past by other authors to study how the operating parameters affect the MB process [10,13]. On the one hand, the quantities that appear on the numerator of equation (3) are the air density at normal conditions ( $ρ_{o} = 1.29 g / L$ , at 0°C and 1 bar) and the air flowrate measured on the flowmeter, corresponding to one set of air exits. The air is dehumidified before it enters the nozzle. On the other hand, the quantities that appear on the denominator are the density of the adhesive ( $ρ_{f} = 0.86 g / {cm}^{3}$ , at 150°C), the adhesive pumped volume per each revolution of the pump ( $v_{c c} = 2.5 {cm}^{3} / rev$ ), the revolutions per minute of the adhesive pump $(n_{rpm})$ , and the number of polymer exits per nozzle $(N_{f})$ . Second, a number that represents the quotient indicating the relative temperature of the air to the fibre will be defined as:

T^{'} = \frac{T_{a} - T_{amb}}{T_{f} - T_{amb}}

(4)

This number is referred to as the dimensionless temperature ratio. This dimensionless parameter is similar to the polymer-air temperature ratio defined by R. Shambaugh [10], but it also includes the ambient temperature $(T_{amb})$ , which was measured in each experiment by means of a thermometer. Moreover, an additional dimensionless number is defined in order to express the distance from the nozzle exit in the $Z$ direction in relative terms and is referred to as the dimensionless distance:

Z^{'} = \frac{Z}{L_{R}}

(5)

Table 2 shows the experimental conditions for the experiments in Table 1 in non-dimensional form.

Table 2.

Operating conditions established for the HSI expressed in dimensionless terms.

A Design			B Design
#Exp.	$m'$	$T'$	#Exp.	$m'$	$T'$
A1	2.60	1.52	B1	4.25	1.52
A2	5.41	1.52	B2	9.01	1.52
A3	1.30	1.52	B3	2.12	1.52
A4	2.60	1.21	B4	4.25	1.21
A5	2.70	1.52	B5	4.51	1.52
A6	5.41	1.21	B6	9.01	1.21
A7	1.30	1.21	B7	2.12	1.21
A8	2.70	1.21	B8	4.51	1.21
A9	2.40	1.34	B9	3.98	1.34

Procedure for measuring the fibre diameter at the nozzle exit region

The formation of shots will be calculated in terms of the coefficient of variation of the fibre diameter at the nozzle exit region. Given the positive results obtained by other researchers when performing automated fibre diameter measurements in melt-blown webs using image processing techniques [14], a tailored MATLAB program was used to measure the fibre diameter. To this end, the image contrast was first adjusted and then it was converted from grayscale into binary format by setting an adequate conversion level. Then, the program split the image into 16 horizontal segments and proceeded to locate the row number (in pixels) of the first row on each of these segments. Then it went through all the pixels of each row and located the first and the last black pixels, which corresponded to the endpoints of the adhesive fibre. Finally, the last step of the measurement procedure of the fibre diameter comprised the subtraction between the pixel column number of these two border-pixels and the subsequent conversion from pixels to millimetres by means of a calibration factor. In order to determine the latter, a millimetre ruler was placed horizontally as a temporary target inside the camera’s FOV, and by using the high-speed cameras software the calibration factor was determined as the number of pixels that correspond to eight millimetres in the ruler. Considerations were also taken on the distortion on the fibre diameter measurement due to bending effects, by multiplying the measurement by a correction factor which depends on the angle that forms the fibre centreline with the horizontal axis. The calculations that lead to the bending correction factor can be found in the Supplemental Material. Figure 6 depicts the procedure followed for measuring the fibre diameter. After measuring the fibre diameter at numerous points below the nozzle exit, the most adequate location for calculating the coefficient of variation had to be selected. In the recorded videos, it was observed that the shots form right at the polymer exit before descending. However, the measurements had to be taken far away enough from the nozzle exit, in a location where the fibre diameter stays constant until it reaches the collector screen. The point along the $Z$ coordinate at which the fibre diameter freezes and no longer varies is referred to as the freezing point. From the diameter measurements, the location of such a point was estimated to be at about $Z^{'} = 12$ on both designs, which is represented in the rightmost picture in Figure 6 by a horizontal dashed line. Hence, at this point the coefficient of variation was calculated on all the experiments.

Figure 6.

Sketch illustrating the measurement procedure of the fibre diameter at the nozzle exit region.

No measurements were performed in cases in which either fibre breakup occurred (experiments B2, B4, B5, B6, B8, B9) or the presence of loops did not allow for precise measurement of the fibre diameter (experiment B7). The results obtained by performing one repetition per each experiment turned out to tally well with the results from the primary experiments, the difference in the fibre diameter at $Z' = 12$ being 49 µm in average.

The uncertainty in the fibre diameter measurements was calculated based on three main components: the uncertainty in the image conversion factor in going from grayscale to binary format, the uncertainty in the calibration factor and the uncertainty due to the variation in time. The uncertainty was calculated following the procedure suggested by R. Moffat [15] and was estimated to be about 11 and 9 µm in the A and B designs, respectively. Details concerning the uncertainty calculations can be found in the Supplemental Material.

Momentum balance

A momentum balance is determined on an adhesive fibre in order to find out how the geometric parameters of the nozzle affect the horizontal and vertical components of the air force exerted on it, which will be ultimately useful for explaining the cohesive fracture of the fibre. The analysis will focus in a set of exits of the nozzle, that is, one polymer exit with its surrounding air exits. For convenience, it is assumed that each air exit has an identical counterpart opposite to the fibre exit. Consider the control volume (CV) shown in Figure 7, with its inner surface delimited by the polymer fibre, its upper surface delimited by the nozzle exit and the side surface defined as the stream surface that delimits all the air jets that correspond to a single polymer exit $(N_{a / f})$ .

Figure 7.

Control volume chosen for analyzing how the momentum is transferred from the air jets to the polymer fibre.

The quotient between the horizontal (along $\hat{Y}$ ) and the vertical (along $\hat{Z}$ ) components of the air force on the polymer fibre represents how much momentum is allocated to squash the fibre when compared with the amount that is allocated to stretch it. As the calculations show (see the Supplemental Material), the force components of the air jets on the polymer fibre along the $\hat{Y}$ and $\hat{Z}$ directions can be expressed by equations (6) and (7):

{(F_{a / f})}_{Y} = \frac{N_{a / f}}{2} ‖ {\vec{\dot{p}}}_{k} ‖ \sin ψ \cos θ

(6)

{(F_{a / f})}_{Z} = \frac{N_{a / f}}{2} ‖ {\vec{\dot{p}}}_{k} ‖ \cos ψ

(7)Where

{\vec{\dot{p}}}_{k}

denotes the momentum flux carried by the air that flows into the CV from a single air exit and

θ

is the precession angle, defined as the angle between

\hat{Y}

and

\hat{ρ}

(see Figure 3). If we divide equation (6) by equation (7), a dimensionless number arises, which is termed the force ratio

(F')

and is given by equation (8):

F_{m h}^{'} = \tan ψ \cos θ

(8)Where the subscript mh stands for multihole. The force ratio of the B design turned out to be

0.62

, while the one of the A design equals

0.13

, being the ratio between both equal to

4.7

. This result reveals that the B Design allocates a higher fraction of the air momentum flux in squashing the polymer fibre than does the A Design, independently of the operating conditions.

The force ratio for an annular nozzle design can be calculated by following a similar procedure, which is also detailed in the Supplemental Material, resulting in equation (9):

F_{a n}^{'} \approx \tan ψ \cos (\frac{7}{25} π)

(9)Where the subscript an stand for annular. Equation (9) shows that an annular nozzle would present a force ratio equivalent to a multihole nozzle that had a precession angle equal to

\frac{7}{25} π

radians (50.4°). As the precession angle of the two multihole nozzle designs studied in the present work is lower than

\frac{7}{25} π

radians, by comparing equations (8) and (9) it can be concluded that both nozzles present a higher force ratio than the annular nozzle design. As will be shown in Fibre breakup section, the force ratio will turn out to be meaningful when comparing the fibre breakup capability of MB nozzles.

Results and discussion

Fibre dynamics

The fibre dynamics were recorded at the nozzle exit region in both nozzle designs using HSI. The fibre exhibited a different dynamical response to the airflow field in each of the two nozzle designs.

In previous work on the classification of the polymer fibre response in different operating conditions, R. Shambaugh differentiated three different regions of melt blown operation [10]. Region I (RI) represents the low gas velocity region where the fibre has little or no oscillating motion and the fibre stays nearly parallel to the air stream, below the air impinging point. If the airflow is increased, the fibre enters Region II (RII), where some adhesive lumps, commonly known as shots, form along the fibre thread. Finally, if the air flow rate continues to increase, the shots are attenuated, and the resulting melt blown fibre is very fine, which represents Region III (RIII). The present results not only confirm the existence of these three regions but also introduce some additional ones that arise from the transition between some regions. The same naming convention introduced by Shambaugh will be used in this work.

In what follows and based on the high-speed camera recordings, we describe the transition that the fibre experiences as the operating conditions vary (only for the operating range presented in Table 1), maintaining and expanding the naming convention introduced by Shambaugh. The description is given from low to high values on the air-polymer flux ratio $(m')$ and from high to low values on the dimensionless temperature ratio $(T')$ . This order corresponds to increasing fibre instabilities. As the value of the first one increases and the second one decreases, the whipping motion and the formation of shots increases in the A design, whereas in the B design the fibre deformation increases and even fibre breakup occurs when $m'$ is sufficiently large and $T'$ is sufficiently low. The regions observed in the A design will be described in first place, while in the second place the ones observed in the B design will be detailed. Details about each region of operation will be provided as the description is given.

For the A nozzle design, the fibre produced experiences Region I and Region II. Region I occur at low $m'$ and high $T'$ , and it consists of a thick stable adhesive fibre which oscillates at a low frequency and presents a low maximum amplitude of oscillation. As $m'$ is increased or $T'$ is decreased, the fibre enters Region II, which is mainly characterized by the accumulation of adhesive near the nozzle tip, and its subsequent fall. As will be shown in Shot section, this polymer accumulation at the nozzle tip will turn out to be responsible for the high variation of the fibre diameter at the freezing point when operating in Region II, representing the main cause of the formation of shots. In Region II, the oscillating frequency and the maximum oscillating amplitude of the fibre increase relative to the values of Region I. From now on, Region II will be called the Shot-Production Region. The two operating regions experienced by the fibre in the A design are depicted in Figure 8(a) and (b) respectively. The geometrical air impinging point is represented by means of a horizontal dashed line and the projection of the air jets on the $Y Z$ plane is represented with two slanted yellow solid lines.

Figure 8.

Melt blowing operating regions present in the A design: (a) Region I; (b) Region II and in the B design; (c) Region I; (d) Region I-III; (e) Region III'; and (f) Region III.

For the B nozzle design, on the other hand, the fibre experiences Region I and Region III but not Region II. However, in this case some subtle observations should be pointed out, especially about the response of the fibre when transitioning from Region I into Region III. First, when operating at a low $m'$ and a high $T'$ , the fibre behaves in a manner similar to that described for the A design in Region I, but with a higher frequency of oscillation. As $m'$ is increased or $T'$ is decreased, the fibre will approximate Region III. However, as the transition between Regions I and III takes place, the fibre will experience one of two transition regions (called Region I-III and Region III’) depending on how $m'$ and $T'$ vary. The dynamical behaviour of the fibre in Region I-III (RI-III) is similar to the one in Region I, the only difference being that occasional fibre breakup occurs at the nozzle exit. But in Region III’ (R III’) the adhesive fibre is flattened into an adhesive membrane, which later re-forms into a single fibre. Unlike the occasional fibre breakup observed in Region I-III, in Region III’ steady fibre breakup occurs. When operating in Region III’ the energy provided by the air jets to the fibre is sufficiently large to deform it and to produce fibre breakup, but it does not suffice to completely overcome the cohesive force of the adhesive, which is apparently responsible of re-forming the adhesive filaments into a single fibre. Finally, the fibre enters the final region, Region III (RIII), which is characterized by the fibre’s complete breakup into several short and thin adhesive filaments, which is commonly known as melt spraying. From this point forward, Region III’ and Region III will be considered the Fibre-Breakup Regions. The operating regions experienced by the fibre in the B design are depicted in Figure 8(c) to (f), respectively.

Figure 9 outlines the fibre dynamical response to the different operating conditions for the two nozzle geometries. For the operating conditions studied in the present work, the A nozzle design does not present a Fibre-Breakup Region and the B nozzle design does not present a Shot-Production Region, thus it can be concluded that both the shot formation and the fibre breakup depend on the nozzle geometry. In agreement with past work on this subject [10], the present work confirms that the shot production does not take place at low air-polymer flux ratios and low polymer temperatures. As the fibre experiences dynamical transitions with increasing values on $m'$ or $T'$ in both figures, it becomes clear that both the production of shots and the onset of the fibre breakup also depend on the operating conditions.

Figure 9.

Melt blowing regions observed in each nozzle design as a function of operating conditions, A design (left), B design (right).

Moreover, in order to relate the different operating regions observed with the high-speed camera at the nozzle exit region with the resulting hot melt application pattern produced by the nozzles, additional images were recorded from the resulting hot-melt applications at the collector screen. The screen was located 10 mm away from the nozzle exit and moved at 100 m/min. These images contain the pattern followed by the polymer fibre at almost identical operating conditions to those presented in Table 2, the air temperature being the only difference between them. While the air temperature $(T_{a})$ in the experiments was always held at 200°C, the images on the screen correspond to two air temperatures, namely 190°C and 210°C. However, given that no significant differences in the hot-melt application pattern has been observed with the different air temperatures, both set of images will be considered to be representative for the adhesive application patterns of the experimental conditions presented in Table 2. The application patterns captured at both temperatures on each region are shown in Figure 10.

Figure 10.

Representative photographs of the hot-melt application patterns obtained from experimental tests, classified by the melt blowing region. The scales that appear on the figure are approximate.

Figure 10(a) and (b) represent regions RI and RII observed in the A Design, and Figure 10(c) to (f) depict regions RI, RI-III, RIII’ and RIII, respectively, observed in the B Design. The upper set of images in Figure 10 corresponds to the experiments conducted with $T_{a} = 190$ °C, while the bottom set corresponds to $T_{a} = 210$ °C. These images confirm the presence of shot (Figure 10(b)) and melt spraying (Figure 10(e) and (f)) on the resulting hot-melt pattern. In addition, from these images it can also be noticed that fibre breakup also occurs in Region I-III to a lesser extent (see Figure 10(d)), which is coherent with the fact that, together with Region III’, these represent regions of transition into Region III.

Fibre breakup

The horizontal and the vertical components of the air force on the polymer fibre were calculated by first substituting the parameters of each nozzle design in equations (6) and (7) and then normalizing them. The normalization was done by subtracting the minimum value obtained for each force component from each force value and then dividing them by the range of variation for each force component and is represented with an overline. The horizontal component of the force ${(\bar{F}}_{Y})$ has been plotted against the vertical one ${(\bar{F}}_{Z})$ in two plots, as shown in Figure 11, according to the value of $T'$ . The data points corresponding to each nozzle design were joined by a straight line.

Figure 11.

Plot of the horizontal force component exerted by the air jets on the fiber against the vertical one as a function of the polymer temperature: (a) $T^{'} = 1.21$ , (b) $T^{'} = 1.52$ .

The slope of the line of the B design to the one of the A design, equals the quotient from dividing the dimensionless force ratio $(F')$ of the first one to the one of the second one (see equation (8)), which results to be proximate to 4.7. Figure 11(a) corresponds to the experiments that were conducted with $T^{'} = 1.21$ , while the one shown in Figure 11(b) corresponds to those with $T^{'} = 1.52$ . Notice that experiments A9 and B9 were conducted at $T^{'} = 1.34$ , and thus these experiments were excluded from Figure 11. As neither equation (6) nor equation (7) includes the polymer mass flow rate $({\dot{m}}_{f})$ , each data label that appears in Figure 11 represents two experiments with different operating conditions, where the label located on the bottom indicates the region observed with the highest ${\dot{m}}_{f}$ and the one on the top indicates the region observed with the lowest one. The experiment number was also added at each data point so Table 2 can be consulted if necessary. From Figure 11, it can be inferred that fibre breakup is increased in nozzle designs that have a higher value of $F'$ (see Momentum balance section). Nozzle designs that do not have a sufficiently large value of $F'$ will not meet the required dynamic conditions for breaking the fibre. From the previous observation and from Figure 8 it can be concluded that cohesive fibre breakup of melt blowing adhesives occurs in the direction perpendicular to the polymer fibre. This observation is in line with the fact that, as fibres are stretched by the aerodynamic force exerted by the air jets, the polymer chains align with the fibre direction, thus presenting a high strength in the axial direction but a low strength in the transverse one [16].

With regard to past research done in the subject, Rao et al. [17], provided a valuable insight into the cohesive fracture in the melt blown process. In particular, these authors stressed that the fibre tensile strength is a function of $T_{f}$ and, in the case of polypropylene, it ranges from 1x10⁹ N/m² at room temperature up to 3x10⁴ N/m² at 180°C. Although these authors studied the thermal behaviour of a different material from the one used in the present study (see Polymer characterization section), the materials are expected to have a similar thermal response. In this sense, the fibre deformation is expected to present a strong dependence on $T_{f}$ . The fact that the cohesive fracture of the fibre is related to the polymer temperature can be confirmed by comparing Figure 11(a) with Figure 11(b), which represent the fibre regions observed in both designs at different polymer temperatures. Given that in the first figure there are two experiments in which the fibre breaks (B6 and B8), while in the second figure there is only one experiment in which fibre breakup occurs (B2), it can be concluded that the cohesive fracture of the fibre in hot-melt applications also depends on $T_{f}$ , as has been observed previously by other authors for similar applications.

From Figures 9 and 11, the following observations can be made about fibre breakup:

It increases with decreasing values of $T'$ and increasing values of $m'$ , being the air-polymer flux ratio the operating parameter that dominates the fibre breakup.

It increases with increasing values of $F'$ .

These conclusions also tally well with the results presented by other authors (4,17). Let us explain each of the aforementioned effects on the fibre deformation from a physical point of view. Firstly, the effect that $T_{f}$ has on the fibre deformation can be explained by the fact that an increase in $T_{f}$ would translate into a decrease in the polymer’s extensional viscosity, which produces a drop in the viscoelastic forces that act along its cross section, and which ultimately control the fibre deformation [18]. The polymer’s elasticity, which is characterized by the relaxation time of the polymer (see Polymer characterization section), also affects fibre deformation [18]. Secondly, from Matsui’s work [19], it is known that an increase in ${\dot{m}}_{a}$ would translate into a higher drag force on the fibre, causing therefore greater deformation, and this explains that a higher ${\dot{m}}_{a}$ increases the fibre breakup. It should be mentioned that if the aerodynamic force is large enough, other flow instabilities can take place as a consequence of the fibre breakup, such as “flies” [20], which are fine fibres that are not collected on the collector screen, or “melt spraying” [20], which represents significant fibre breakup. Thirdly, the relationship between the fibre deformation and $F'$ is the most difficult one to clarify, because it involves the adhesive’s composition and the polymer’s mass weight distribution, both of which determine how the fibre reacts under the application of external forces. Even the oxidative degradation the polymer fibre undergoes when exposed to the air jets, which produces a stratification in the fibre’s cross section, can affect fibre breakup [21]. Nevertheless, as the fibre is squeezed, it exhibits both extensional and shear deformation, both of which are limited by the polymers molecular-weight [5]. Hence, it is expected that adhesives that have a higher molecular weight than the one studied may not exhibit fibre breakup at the operating conditions of Region (RIII). Hence, given the aforementioned statements, the first two observations do not depend on the polymer being used, whereas the third one is valid only for the adhesive used in the present work, which is a typical commercial polymer used for producing hot melt adhesives for use in the hygiene industry.

These observations suggest that both the operating conditions and the geometric parameters of the nozzle affect fibre breakup. With regard to the first factor, fibre breakup would require operating with high air and low polymer flow rates as well as high polymer temperatures, while the second factor would require designing a MB nozzle with a high global tilt angle. The last observation supports the fact that the annular nozzle design, which presents a lower value of $F'$ than the multihole A and B nozzle designs do, presents a lower potential to break the fibre than the multihole design when working in the same operating conditions. Furthermore, when studying the fibre breakup capability of MB nozzle designs other than the ones treated in the present study, an equivalent definition of the global tilt angle $(ψ)$ should be adopted for calculating $F'$ , especially for those designs that present air ducts with either unequal tilt angles or exit areas.

Shot

The formation of shots was quantified on the A design by means of the coefficient of variation $(c_{v})$ of the fibre diameter $(d_{f})$ at the freezing point $(Z^{'} = 12)$ , which is a measure of the variation of the diameter with respect to its mean value. This parameter is known to vary with the composition of the polymer [22], and the present study analyses its variation with the operating conditions and with the nozzle geometry. If the fibre diameter at the nozzle exit region is assumed to follow a normal distribution, then its coefficient of variation can be calculated as:

c_{v} = \frac{σ_{d_{f}}}{µ_{d_{f}}}

(10)where

σ_{d f}

and

µ_{d f}

are the standard deviation and mean value of the fibre diameter, which will be estimated by means of the sample standard deviation and mean, respectively. The results obtained for

c_{v}

in experiments A1 to A9 are shown in Table 3. From Table 3 it can be inferred that high values of

c_{v}

indicate the formation of shots.

Table 3.

Coefficient of variation of the fibre diameter calculated at the freezing point, normalized die drool area (NDDA) and its standard deviation calculated at $Z' = 1$ .

#Exp.	Region	$σ_{NDDA}$	$c_{v}$	$NDDA$
A1	RI	0.16	0.10	2.19
A2	RII	0.29	0.81	1.24
A3	RI	0.09	0.12	2.21
A4	RII	0.44	0.68	1.84
A5	RII	0.24	0.86	1.40
A6	RII	0.25	0.70	1.24
A7	RII	0.18	0.45	2.13
A8	RII	0.26	0.64	1.37
A9	RII	0.30	0.41	1.56

Boldface indicates Shot-Production Region.

After analysing the videos recorded with the high-speed camera, it was observed that the shot was produced under certain operating conditions and after a critical mass of polymer had accumulated at the nozzle tip, forming an adhesive lump which subsequently descended and caused large fibre oscillations. The accumulation of extrudate at the nozzle tip is a phenomenon termed as die drool [23]. There are two distinct mechanisms that lead to die drool. The first one takes place inside the adhesive capillary (internal) and the second one takes place at the nozzle exit region (external). Whereas negative pressure gradients at the nozzle exit region can lead to external die drool, internal die drool is caused by molecular-weight fractionation of the polymer [23]. Die drool (or die build-up) can be quantified by measuring the Normalized Die Drool Area (NDDA) by use of Digital Image Analysis (DIA) [23,24] and is defined as:

NDDA = \frac{A_{D}}{A_{Die}}

(11)where

A_{D} = \frac{π d_{f, 1}^{2}}{4}

is the area of the accumulated die drool material at

Z^{'} = 1

and

A_{Die} = \frac{π L_{R f}^{2}}{4}

is the polymer exit area,

d_{f, 1}

and

L_{R f}

being the fibre diameter at

Z^{'} = 1

and the polymer exit hydraulic diameter, respectively. The high-speed images revealed that the adhesive accumulates uniformly around the fibre at the nozzle exit, therefore the cross section of the fibre is assumed to maintain its circular shape at

Z^{'} = 1

, as the definition that has been given for

A_{D}

confirms. Other authors that have observed die drool at the nozzle exit reported that the material can accumulate in structures different from the circular one, such as the torus-shape structure [24], or even in flake or powder form [25]; depending on the sample being studied and the process conditions.

NDDA

was calculated at

Z' = 1

in all the experiments and the results are shown in Table 3. From Table 3 it can be observed that the standard deviation of

NDDA {(σ}_{NDDA})

, which has also been calculated and added to Table 3, maintains a positive correlation with

c_{v}

. This observation shows that the variation of accumulated mass at the nozzle exit is the root cause of shot formation. In other words, the mass that accumulates at the nozzle exit sticks to the extrudate and descends to form fibre shots on the adhesive pattern. Hence, a higher

c_{v}

or, alternatively, a higher

σ_{NDDA}

leads to larger shot size and more frequent shot formation.

The descent of the accumulated mass at the nozzle exit causes parameter $NDDA$ to oscillate with time, instead of exhibiting an exponential growth, as it does in nozzles on which the mass accumulates steadily at the nozzle tip. Figure 12 shows four plots of $NDDA$ as a function of time for experiments A1 to A8. The plots have been grouped by the value of the air-polymer flux ratio. The dotted lines correspond to experiments performed at $T^{'} = 1.21$ (higher $T_{f}$ ), while the solid lines correspond to experiments performed at $T^{'} = 1.52$ (lower $T_{f}$ ). On the one hand, the formation of shots with decreasing values on $T'$ (increasing values on $T_{f}$ ) can be explained by the increase in the polymer’s molecular weight fractionation, stress-induced cohesive chain disentanglement and polymer decomposition, all of which favour the formation of a hydrodynamic layer close to the nozzle walls, which increases internal die drool [23]. As a result, no sooner is the adhesive ejected from the nozzle than the external low-molecular layer of adhesive builds up at the nozzle exit. On the other hand, the formation of shots with increasing values of $m'$ can be explained by the increase in the pressure drop at the nozzle exit (caused by the increase in air velocity), which increases external die drool. In addition, nozzle designs with smaller global tilt angles $(ψ)$ will present a larger low-pressure region at the nozzle exit, which ultimately favours die drool and consequently the formation of shots. This observation is coherent with the fact that $ψ$ in the A design is 5.6 times smaller than the one in the B design (see Description of the nozzle geometry section), and shots were observed only in the first one. Figure 9 not only confirms that the formation of shots increases with increasing $m'$ and decreasing $T'$ , but also shows that $m'$ is the operating parameter that most influences the formation of shots. The previous explanation that was given for shot formation confirms the hypothesis that was proposed by Breese and Yan with the direct shot formation model, which suggests that shots form at the nozzle exit.

Figure 12.

Normalized Die Drool Area as a function of time: (a) $m^{'} = 1.30$ , (b) $m^{'} = 2.60$ , (c) $m^{'} = 2.70$ , (d) $m^{'} = 5.41$ .

The previous observations are in agreement with past research on shot formation, which establishes that the polymer’s molecular weight, which is a measure of internal die drool, as well as the air tilt angle, which is a measure of external die drool, are some of the main parameters that explain this phenomenon [26]. Regarding the formation of shots and internal die drool, morphological differences in the internal structure of fibre strands and shot particles have been observed by use of scanning electron microscopy [9]. However, to the author’s knowledge, no study has reported results on the molecular weight distribution of shot particles. Jiri Drabek and Martin Zatloukal [22] determined the fibre diameter’s coefficient of variation $(c_{v})$ of linear PP and branched PP (both having comparable molecular weight and polydispersity) at $DCD = 200$ mm and $DCD = 500$ mm and observed that linear polymer increases the variation of the fibre diameter, which can also be explained as a result of an increase in the molecular fractionation [27,28]. However, the authors do not relate such variation to the formation of shots.

Our studies suggest that the measures that show to be effective in reducing die drool will also help to minimize the formation of shots. Some of these include reducing molecular fractionation, minimizing material degradation, reducing the contents of lubricants or waxes and supplying clean hot air around of the extrusion orifice [23].

Figure 13 shows a representative sequence of frames taken in experiment A2 (RII), in which the shot formation was captured. This sequence of frames illustrates not only the shot formation but also the descent process through eight milliseconds of recording time, where the time shift between two consecutive frames is of about one millisecond. In this figure it can be observed how the shot descends across the freezing point $(Z^{'} = 12)$ , which is represented by a horizontal dotted line, showing no signs of attenuation. The location at which $NDDA$ was quantified $(Z^{'} = 1)$ is represented by a horizontal dashed line. The presence of shots within the hot melt application leads to an increase in the variation in the fibre diameter, jeopardizing the uniformity of the adhesive bond.

Figure 13.

Sequence of frames presenting the shot formation process, taken from experiment A2, with a time shift between frames of 1 ms.

Although there are operating conditions under which shots do not form, such as in experiment A3, and there are yet others under which there is enough evidence to show that they do form, such as in experiment A4, there are nevertheless some operating conditions under which shots form to a lesser extent, as some other experiments show. One of these examples is shown in Figure 14, which depicts the adhesive pattern obtained in experiment A7, for which the calculated $c_{v}$ is 0.45 (see Table 3). From this observation, the threshold $c_{v}$ that leads to the formation of shots is estimated to be about 0.45.

Figure 14.

Adhesive pattern obtained for experiment A7, which corresponds to $c_{v} = 0.45$ .

The results obtained in the previous and present sections indicate that there is a compromise between the geometric parameters of the nozzle and the operating conditions, both of which determine fibre breakup and shot formation. Finally, it should be mentioned that the conclusions that emerge from the present work on shot formation will be valid only for the adhesive being studied.

Concluding remarks

In this article, fibre deformation and shot formation were studied in the melt blowing process at the nozzle exit region on two multihole nozzle designs from an experimental point of view. To this end, the fibre dynamics were captured by using HSI for various operating conditions, which were taken from real industrial hot-melt adhesive applications for hygiene products.

On the one hand, it was shown that fibre breakup occurs as a result of applying a sufficiently large force in the direction perpendicular to the fibre. On the other hand, it was shown that the root cause of fibre shots is the accumulation of mass at the nozzle exit region (die drool), which consequently increase the variation in the final fibre diameter. Fibre shot can be quantified by means of the coefficient of variation of the fibre diameter $(c_{v})$ at the freezing point and the maximum $c_{v}$ without leading to shots in the adhesive pattern was estimated to be about 0.45. It is expected that hot-melt adhesives with a higher molecular weight than the one studied will reduce both fibre breakup and shot formation.

The results suggest that both the nozzle geometry and the operating conditions determine whether fibre breakup will occur and whether shots will be produced. Concerning the operating conditions, we concluded that increasing values of either the air-polymer flux ratio or the polymer temperature contributed to breaking the fibre and also facilitated the formation of shots. Concerning nozzle geometry, the results indicate that increasing values on the air duct tilt angle $(ψ)$ has two effects. Firstly, it increases the ratio between the horizontal and vertical force components of the air jets, which in turn increases the fibre’s cohesive fracture. This ratio was shown to be higher in symmetrical multihole nozzles than in annular nozzles. Secondly, it reduces the low-pressure area at the polymer exit, what contributes to reducing external die drool and therefore the formation of shots.

Further study of fibre breakup and shot formation in the melt blowing of adhesives with different compositions other the one used in the present work or on different nozzle designs would provide a better understanding of these two defects, which would ultimately permit nozzle technology to be improved for use in the blown fibre production.

Supplemental Material

sj-pdf-1-jit-10.1177_1528083720949276 - Supplemental material for Experimental study of fibre breakup and shot formation in melt blowing nozzle designs

Supplemental material, sj-pdf-1-jit-10.1177_1528083720949276 for Experimental study of fibre breakup and shot formation in melt blowing nozzle designs by Ignacio Formoso, Alejandro Rivas, Gerardo Beltrame, Gorka S Larraona, Juan Carlos Ramos, Raúl Antón and Alaine Salterain in Journal of Industrial Textiles

Footnotes

Acknowledgements

The authors are grateful for the support of Cátedra Fundación Antonio Aranzábal-Universidad de Navarra. The authors also want to acknowledge Juan Villarón for his immense help in making the HSI recordings.

Declaration of conflicting interests

The author(s) declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The author(s) disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: This work was supported by VALCO MELTON; and Ministerio de Ciencia e Innovación, Gobierno de España, RETOS-COLABORACION 2019 (RTC2019-007057-7).

ORCID iD

Ignacio Formoso

Supplemental Material

Supplemental material for this article is available online.

References

Gahan

Zguris

GC.

A review of the melt blown process. In: Proceedings of the Annual Battery Conference on Applications and Advances. Piscataway: IEEE, 2000, pp. 145–149.

Allen

Fetcko

JT.

Modular meltblowing die. Patent Number: 5618566, 1997.

Hao

Zeng

A review on the studies of air flow field and fiber formation process during melt blowing. Ind Eng Chem Res 2019; 58: 11624–11637.

Ellison

Phatak

Giles

, et al. Melt blown nanofibers: fiber diameter distributions and onset of fiber breakup. Polymer 2007; 48: 3306–3316.

Wang

Boukany

Wang

, et al. Elastic breakup in uniaxial extension of entangled polymer melts. Phys Rev Lett 2007; 99: 1–4.

Roland

Archer

Mott

, et al. Determining rouse relaxation times from the dynamic modulus of entangled polymers. J Rheol 2004; 48: 395–403.

Han

Bhat

Wang

Investigation of nanofiber breakup in the melt-blowing process. Ind Eng Chem Res 2016; 55: 3150–3156.

Taylor

Yan

Bresee

RR.

Characterizing nonwoven-web structure by using image-analysis techniques. Part V: analysis of shot in meltblown webs. J Text Inst 1998; 89: 320–341.

Bresee

Yan

Shot development in meltblown webs. J Text Inst 1998; 89: 304–319.

10.

Shambaugh

A macroscopic view of the melt-blowing process for producing microfibers. Ind Eng Chem Res 1988; 27: 2363–2372.

11.

Aliani G, Lechat JB, Smits JA, et al. Ethylene copolymers for hot melt systems. Patent 4,613,632, USA, 1985.

12.

McConnell

Meyer

Petke

, et al. Polyester adhesives in nonwovens and other textile applications. J Ind Text 1987; 16: 199–208.

13.

Shambaugh

Kayser

The manufacture of continuous polymeric filaments by the melt-blowing process. Polym Eng Sci 1990; 30: 1237–1251.

14.

Zhang

Wang

, et al. Automated measurements of fiber diameters in melt-blown nonwovens. J Ind Text 2014; 43: 593–605.

15.

Moffat

Describing the uncertainties in experimental results. Exp Therm Fluid Sci 1988; 1: 3–17.

16.

Ashby

MF.

Materiales para ingeniería 2. Barcelona: Editorial Reverté SA, 2009, pp. 319–320.

17.

Rao

Shambaugh

Vibration and stability in the melt blowing process. Ind Eng Chem Res 1993; 32: 3100–3111.

18.

Tan

Zhou

Ellison

, et al. Meltblown fibers: influence of viscosity and elasticity on diameter distribution. J Non-Newton Fluid Mech 2010; 165: 892–900.

19.

Matsui

Air drag on a continuous filament in melt spinning. Trans Soc Rheol 1976; 20: 465–473.

20.

Drabek

Zatloukal

Meltblown technology for production of polymeric microfibers/nanofibers: a review.

Phys Fluids 2019; 31: 091301.

21.

Kelley

Shambaugh

Sheath-core differences caused by rapid thermoxidation during melt blowing of fibers. Ind Eng Chem Res 1998; 37: 1140–1153.

22.

Drabek

Zatloukal

Influence of long chain branching on fiber diameter distribution for polypropylene nonwovens produced by melt blown process. J Rheol 2019; 63: 519–531.

23.

Musil

Zatloukal

Historical review of die drool phenomenon in plastics extrusion. Polym Rev 2014; 54: 139–184.

24.

Musil

Zatloukal

Investigation of die drool phenomenon for HDPE polymer melt. Chem Eng Sci 2010; 65: 6128–6133.

25.

Chaloupková

Zatloukal

Theoretical and experimental analysis of the die drool phenomenon for metallocene LLDPE. Polym Eng Sci 2007; 47: 871–881.

26.

Goswami

BC.

Spunbonding and melt-blowing processes. Dordrecht: Springer, 1997, pp. 560–594.

27.

Inn

Melt fracture and wall slip of metallocene-catalyzed bimodal polyethylenes in capillary flow. J Rheol 2013; 57: 393–406.

28.

Musil

Zatloukal

Experimental investigation of flow induced molecular weight fractionation during extrusion of HDPE polymer melts. Chem Eng Sci 2011; 66: 4814–4823.

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