A prototype X/S-band common aperture horn feed for future use at various DSN sites and the Network Consolidation Program is discussed. The final design and fabrication of the second generation feedhorn and combiner is dealt with. The results of the measurements obtained with the second generation, full scale feed configuration are presented

Reilly, H.

Williams, W.

English

NASA Technical Reports Server

TDA Progress Report 42-60 September and October 1980 
A Prototype DSN XIS-Band Feed: DSS 13 Application Status 
(Fourth Report) 
W. Williams and H. Reilly 
Radio Frequency and Microwave Subsystems Section 
This article, the fourth in this series, discussing a prototype XIS-band common 
aperture hom feed for future use at various DSN sites and the Network Consolidation 
Program, deals with the final design and fabrication of the second-generation feedhom 
and combiner. The results of the measurements obtained with the second-generation, 
full-scale feed configuration are presented. 
I. Introduction 
The first three articles (Refs. I, 2, and 3) in this series 
discussed the development of a first-generation, dual-band 
(X-S) corrugated feedhorn for DSN applications. The earlier 
articles presented horn radiation patterns and resultant reflec-
tor antenna efficiencies using a half-scale horn model, the 
design of an X-S combiner permitting injection of both bands 
into the horn, and finally the measurements on the full-scale, 
first-generation horn and combiner. 
The first horn model tests gave evidence ofa small amount 
of "moding" in the X-band, but still within the specification 
requirement, and the X-S combiner had just enough S-band 
frequency bandwidth to satisfy the original narrow S-band 
requirement of approximately 40 MHz. 
A second-generation horn system was begun and evolved 
where the major requirement was to increase the S-band per-
formance across sufficient range to encompass DSN S-band 
high-power transmission as well as the DSN receive bands (2.10 
to 2.30 GHz). As an adjunct, it was desired to determine the 
cause of X-band moding and, if possible, to correct it. The 
technique for broadening the combiner S-band bandwidth was 
by increasing the combiner radial line height and carefully 
designing a wideband matching network; the method of cor-
recting X-band moding is discussed in Ref. 3. 
II. The Full-Scale, Second-Generation XIS 
Horn 
Early investigations into the causes of X-band moding indi-
cated that the abrupt horn input angular change from cylindri-
cal waveguide to the 34-degree wide-angle horn caused some 
moding, but probably not as much as noted. Therefore, the 
final horn version was to be built using a gradual change to the 
34-degree flare angle over a 100-mm transition length to par-
tially reduce moding. 
A unique concept for this corrugated horn, relative to 
others that have been used, is the abrupt change in corrugation 
(groove) depth in the region of the horn where S-band is 
introduced. Grooves must be between %- and ~·wavelength 
deep to generate the proper corrugated waveguide hybrid 
mode. When cut the required depth for S-band, they become 
from I %- to I ~-wavelength deep in X-band, and hence will 
generate the proper X-band boundary conditions and modes. 
77 
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However, the particular configuration of the combiner design 
in this program did not allow the full groove depths to be used 
in the horn input region where only X-band was present and so 
shallower grooves were used (~ to % wavelength at X-band) at 
the input. This then necessitated an abrupt change near the 
combiner S-band input from the ~-wavelength groove 
(X-band) to the 1 ~ wavelength groove (X-band). This change 
represents a potential discontinuity in groove impedance for 
the X-band, dependent upon X-band frequency. 
Experiments with sections of the first-generation horn 
revealed that this impedance discontinuity was the major cause 
of extraneous moding, (indicated by the level of cross polariza-
tion in the 4S-degree pattern cut). Therefore, groove depths 
were cut that gave a groove impedance match near the center 
of the desired X-band range, and hence symmetrical mis-
matches at the band edges; e.g., at 7.8 GHz, the two groove 
depths at the abrupt change were 0.297A (A = X-band wave-
lengths) and 1.297A, giving a "perfect" match. This resulted in 
measured -43-dB cross polarization, a good result indeed. At 
7.1 GHz, the selected depth results in a 0.271 A input groove 
and a 1.l8A output groove, a mismatch giving rise to a cross-
polarization level of -28 dB; however, this is still acceptable. 
At 8.45 GHz, the selected (7.8-GHz) depth results in a mis-
match from 0.322A to 1.40SA, and similar (-28-dB) cross-
polarization levels. 
Therefore, the full-scale, second-generation horn was made 
with X-band input grooves at 11.43 mm and output grooves at 
49.89 mm, giving the ideal match at midband, 7.8 GHz, unfor-
tunately a frequency region of no planned use. Our intentions 
are to provide 7.l-GHz and 8.4-GHz bands in the second-
generation horn. 
In the first-generation horn, the grooves were but 3.5S-mm 
wide, the object being to have as many grooves per X-band 
wavelength as possible. This has always been reasonable in the 
past because such an assumption was made in theoretically 
solving for the hybrid mode fields in corrugated waveguide. 
However, in our second-generation horn (at the suggestion of 
B. M. Thomas, Ref. 4) the grooves were cut much wider; the 
only requirement being that they remain less than %-wave-
length wide at the highest in tended frequency. This has 
worked quite well and permitted an easier (higher waveguide) 
access for the S-band horn input. The second generation horn 
has grooves of 12.7-mm width with a 3.SS-mm wall. 
The full-scale, second-generation horn was machined from 
five billets of aluminum with the smaller X-band input section 
and the XIS combiner made from separate pieces. The first 
generation horn exterior photograph given in Ref. 3 is not 
unlike our later development. 
78 
The radiation patterns of the second-generation horn have 
been measured at many frequencies throughout both the X 
and S bands. These measurements are made using linear polar-
izations. At each frequency, an E- and an H-plane pattern were 
recorded and also a pattern in the 4S-degree plane. Also in the 
4S-degree plane the cross-polarization pattern was recorded, 
this indicating the lack or presence of unwanted radiating 
modes. E- and H-plane patterns are presented together on one 
graph while a second graph shows the 4S-degree plane matched 
and cross polarized. 
Figure 1 presents the results at X-band while Fig. 2 presents 
results at S-band. The E- and H-plane equivalence is very good 
and the complete lack of sidelobes down to minus 40 dB (and 
lower) contributes to a uniquely higher forward spillover effi-
ciency than has been available before in X-band. Although the 
S-band patterns also are devoid of sidelobes, the shape of these 
patterns results in greater than typical forward spillover. Note 
the relative steepness on the sides of the X-band patterns 
compared to the sides of S-band patterns - a result of main-
taining a sensible horn aperture size. 
III. The Full-Scale, Second-Generation XIS 
Combiner 
The first-generation XIS-band combiner performed its func-
tion well; it extracted the S-band receive signal from the horn 
at low loss, contributing little to system noise performance 
and its X-band isolation was great enough such that any 
additional X-band noise contribution was not detectable. How-
ever, this unit was, at S-band, of such narrow bandwidth that 
it could not be used for simultaneous S-band reception and 
transmission. Broadening the S-band bandwidth of this com-
biner so that receive/transmit functions could be included was 
the major objective of the second-generation development. 
The first-generation combiner is comprised (see Ref. 2 for 
photos and a detailed discussion) of a thin radial line, only 
8.89-mm wide, surrounding the horn at a horn diameter of 
about 119.4 mm and used to inject S-band into the horn. 
Internal to the radial line are a pair of X-band rejection chokes 
that isolated the X-band successfully so that no additional 
X-band noise could be detected as coming from the S-band 
portions. The very narrow radial line was chosen to help assure 
that no X-band energy could penetrate the combiner. How-
ever, this narrow line also tends to make the S-band passband 
more limited (as determined by the impedanc~ variation with 
frequency looking into the combiner). The horn entry point 
(119.4-mm diam.) was selected for a best impedance match at 
2.3 GHz and this resulted in a bandwidth of approximately 
50 MHz over which input VSWR was less than 1.2:1. 
In the second-generation combiner, the radial line section 
has been increased to 12_7 mm to increase the S-band band-
width performance to include S-band transmission from 2.1 
to 2.3 GHz. The band center is now lower (2.2 GHz) and 
therefore a new injection horn diameter of 127 mm is used. 
An increase to four X-band reject radial line chokes is used to 
maintain the required high X-band isolation. As with the 
first-generation combiner, tuning irises were required to be 
inserted into the radial line area to achieve an acceptable 
performance across the S-band. It was convenient to place 
these irises as metal blocks inserted into portions of the 
X-band chokes. A disassembled picture of this combiner is 
shown in Fig. 3. 
The combiner along with its four 12.7 -mm input waveguide 
terminals is not matched to standard S-band waveguide. The 
additional matching must be done with a transformer-tuner 
that transforms the 12.7-mm waveguide height to the 54.6-mm 
height of standard S-band (WR430) waveguide and at the 
same time provides the necessary tuning to match the input 
impedance over the required bandwidth to a VSWR (voltage 
standing wave ratio) of less than 1.2:1. The graph shown in 
Fig. 4 represents the VSWR looking into anyone of the four 
combiner input terminals. The requirement on the tuner is to 
develop a response VSWR of less than 1.2: 1 from 2.1 to 
2.3 GHz. 
The transformer-tuner was designed and is fabricated using a 
3-step, 2-section waveguide section to transform from the 
12.7-mm height waveguide combiner input (standard 109 .2-mm 
width), down to a narrow height guide of only 3.12 mm, 
increasing to a section of 39.4-mm height and then to the full 
54.6 mm of standard waveguide. At the final step, an inductive 
iris is inserted into the 39.4-mm size that tunes the total 
combination across the required band. The VSWR response of 
the transformer-tuner and combiner is also shown in Fig. 4. 
Note that there exists a small region above 2.25 GHz where 
the VSWR exceeds 1.2: 1, but this receive-only region is less 
critical than the lower or transmit end of the band, and fully 
acceptable. 
The tuner-combiner is designed to transmit 20-kW CW 
power, or 5 kW into each tuner port under circular polariza-
tion excitation. Calculations of voltage breakdown in the 
3.12-mm height section of waveguide indicate that it has this 
capability with sufficient safety factor. However, if the trans-
mitter requirement should become 100 kW or more at a later 
date, there is serious question about the performance of this 
particular transformer-tuner design. 
IV. A Tuner Design for High Power 
Another model transformer-tuner was designed for the 
eventuality of higher power transmission. This unit has but 
two steps and one section to transform the required imped-
ances. The minimum waveguide height of this unit is 12.3 mm 
and will withstand a 200-kW transmitter with sufficient safety 
factor. However, the unit does not meet the 1.2:1 VSWR 
specification across the entire 2.1- to 2.3-GHz S-band, but 
only in two narrow bands that include the present DSN 
transmit (211 0- to 2120-MHz) and receive (2290- to 
2300-MHz) bands. The VSWR response of the high-power 
tuner and combiner is also shown in Fig. 4. From Fig. 4, one 
notes that the two bands remain less than VSWR = 1.2 over 
roughly equal 35-MHz bandwidths. 
V. Calculated Performance for DSS 13 
Demonstration Application 
Referring to Ref. 2, p. 41, one notes the discussion con-
cerning the DSS 13 subreflector having been designed for 
S-band. Again, as in the first generation case, the measured 
horn patterns of Figs. 1 and 2 are used in a scattering program· 
with the DSS 13 subreflector (including the vertex plate and 
outer flange) to determine the final primary reflector excita-
tion and efficiency for the 26-meter paraboloid at DSS 13. 
Two techniques are used to determine these subreflector 
scattered patterns. In one, the measured far field pattern is 
used to determine current excitations on the subreflector and 
finally the physical optics scattering. In the other technique, 
the horn radiation pattern is used to determine its spherical 
wave coefficients (Ref. 5) and these are then used to deter-
mine currents on the subreflector at its unique range from the 
horn, instead of assuming far field. These techniques agreed to 
within 0.5 percent, and so the far field approach is used for all 
calculations herein. 
Figure 5 shows the DSS 13 scattered patterns in X-band 
and S-band. One can observe the effect of the S-band designed 
vertex plate at X-band with a smaller, perhaps modest, effect 
at S-band. Although subreflector blockage of radiated power is 
reduced to essentially zero at X-band, a corresponding X-band 
loss is noted (relative to S-band) in illumination efficiency and 
phase efficiency, due to pattern distortions related to the 
(oversized) vertex plate. 
The efficiencies calculated from these scatter patterns are 
tabulated in Table 1. The 71.3-percent value at X-band is 
about 5 percent higher than the heretofore standard (22-dB) 
horn feed used, or about +0.3 dB. 
79 
Figure 6 presents the calculated far-field patterns from the 
26-meter paraboloid. These patterns neglect the effect of spar 
blocking and surface tolerance; hence all fmal measured 
patterns will have somewhat higher sidelobes. 
The final expected performance at DSS 13 may now be 
obtained as shown in Table 2. These overall efficiencies neglect 
all feed internal losses and final feed VSWR. 
Using the geometrical shadow from the spars at DSS 13 and 
an empirically developed radio-frequency factor for spar 
blockage, a spar blockage efficiency of 0.861 has been deter-
mined. Also the surface tolerance (e) at DSS 13 is less than 
l.5-mm rms. When applied to the surface tolerance efficiency 
formula 
VI. Conclusions 
The XIS second-generation feed system will develop about 
3-percent more gain than the standard 22-dB horn in the DSS 
13 configuration. Also, since the noise contribution from rear 
spillover is less, a slight improvement in system noise tempera-
ture may be expected. A significant advantage will be seen in 
the application of this feed in other configurations, particu-
larly the planned Network Consolidation Project. This will be 
the subject of a later article. (
- {41Te F) 
1) (surface) = exp A 
there results the surface tolerance efficiencies The concluding work on the horn system will take place at 
the DSN Microwave Test Facility where system noise and 
power capability will be measured. This will be the subject of a 
final article. 
80 
75.4 percent at 8.450 GHz 
97.9 percent at 2.295 GHz 
References 
l. Williams, W. F., "A Prototype DSN X-S Band Feed: DSS 13 First Status Application," 
DSN Progress Report 42-44, January and February 1978. Jet Propulsion Laboratory, 
Pasadena, Calif. 
2. Williams, W. F., "A Prototype DSN X-S Band Feed DSS 13 Application Status 
(Second Report)," DSN Progress Report 42-47, July and August 1978. Jet Propulsion 
Laboratory, Pasadena, Calif. 
3. Williams, W., et aI., "A Prototype DSN XIS Band Feed: DSS 13 Application Status 
(Third Report)," DSN Progress Report 42-52, May and June 1979. Jet Propulsion 
Laboratory, Pasadena, Calif. 
4. Thomas, B. Mac A., "Design of Corrugated Horns," IEEE Trans. on Ant. and Prop., 
Vol. AP-26, No.2, March 1978. 
5. Ludwig, Arthur C., "Near-Field Far-Field Transformations Using Spherical Wave 
Expansions," IEEE Trans. on Ant. and Prop., Vol. AP-19, March 1971, pp. 214-220. 
Table 1. Second-generation DSS 13 efficiencies 
Efficiency Frequency = Frequency = 8.450 GHz 2.290 GHz . 
Rear spillover 0.997 0.994 
Forward spillover 0.979 0.889 
Illumination 0.811 0.865 
Cross-polarization 0.999 0.999 
Phase 0.902 0.925 
Blockage (subreflector) 1.0 0.972 
Total 0.713 0.686 
Table 2. Second-generation DSS 13 final overall efficiency 
Efficiency Frequency = Frequency = 8.450 GHz 2.295 GHz 
RF feed efficiency 0.713 0.686 
Surface efficiency 0.754 0.979 
Spar blockage 0.861 0.861 
Final overall efficiency 0.463 0.578 
81 
82 
0 (a) (b) 
4 
E-PLANE 
8 
12 
16 
20 
24 CROSS 
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I 
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12 
16 
20 
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36 
72 36 36 n n 36 0 72 
ANGLE, deg 
Fig. 1. Radiation patterns of the second-generation feedhorn in X-band: (a) E and H plane patterns at 8.45 GHz; (b) pattern cut at 45 deg to 
show cross polarization at 8.45 GHz; (c) E and H plane patterns at 7.8 GHz; (d) pattern cut at 45 deg to show cross polarization at 7.8 GHz; 
(e) E and H plane patterns at 7.1 GHz; (f) pattern cut at 45 deg to show cross polarization at 7.1 GHz 
0 (e) (0 
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Fig. 1 (contd) 
83 
84 
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8 
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12 
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ANGLE, deg 
Fig. 2. Radiation patterns of the second-generation feedhorn in S-band: (a) E and H plane patterns at 2.3 GHz; (b) pattern cut at 45 deg to 
show cross polarization at 2.3 GHz; (c) E and H plane patterns at 2.1 GHz; (d) pattern cut at 45 deg to show cross polarization at 2.1 GHz 
Fig. 3. The Mod II XIS Combiner 
85 
86 
3.5r-----------------------------------~ 
3.0 
2.5 
1.5 
1.0L-____ ~L_ __ _L ____________ ~ ____ ~~ __ ~ 
2050 2150 2250 2350 
FREQUENCY, MHz 
Fig. 4. The VSWR at input to the second-generation combiner and at 
the tuner input 
o'(-a-)-'---'-d-E-TH'E-TA---'---'-I--'I-~-~-'I----~--TI--hh~.II---'I----'-1--'---'----'---'-1--'---' 
-5- r ~ft~ -X E PHI • • l~' 
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-15 
-20 
-25 
-30 
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~ (b) 
~ 
~ -5 
-10 
-15 
-20 
POLAR ANGLE, deg 
Fig. 5. Subreflector scattering at azimuth angle = 0.00, second-generation horn from the DSS 13 subreflector: (a) at 8.450 GHz; 
(b) at 2.295 GHz 
87 
-5 
-10 
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88 
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POLAR ANGLE, deg 
Fig.6. Secondary patterns at azimuth angle = 0.00 from the Venus paraboloid (DSS 13): (a) at 8.450 GHz; (b) at 2.295 GHz 


A prototype DSN X/S-band feed: DSS 13 application status

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